Brijesh Singh, Vikas Patel, P N Ojha, VV Arora · 2020-07-16 · Brijesh Singh, Vikas Patel, P N Ojha, VV Arora ... [1,2]. However, high-strength concrete is more brittle in nature

Journal of Asian Concrete Federation

Vol. 6, No. 1, pp. 1-9, June 2020

ISSN 2465-7964 / eISSN 2465-7972

https://doi.org/10.18702/acf.2020.6.6.1

1

Technical Paper

Analysis of stress block parameters for high strength con-

crete

Brijesh Singh, Vikas Patel*, P N Ojha, VV Arora

(Received January 27, 2020; Revised June 10, 2020; Accepted June 21, 2020; Published June 30, 2020)

Abstract: In Indian Standard code IS: 456-2000 for concrete of grades higher than M55, the design

parameters given may not be applicable as structural behaviour of concrete changes as strength of con-

crete increases. Different international standards give different stress block parameters which can be re-

duced to these two basic factors. In this paper, stress block parameters K (strength reduction factor) and

k2 (factor for the depth of resultant compressive force) were calculated from experimental strain values

considering the stress-strain curve as parabolic for lower grades and as linear for higher grades. Also, the

method and approach for the calculation of stress block parameters have been worked out. The method

so proposed is compared with the model proposed in European design standard EC: 02-2004 by trans-

forming the European stress block parameters to the basic parameters used. Also, the effect of the shape

of the stress-strain curve and value of ultimate strain in concrete on stress block parameters and moment

capacity of the members was analyzed by working on the representative section. The method or approach

so proposed will be useful to understand and compare flexural design philosophies used in different in-

ternational standards by reducing the stress block parameters to two basic factors.

Keywords: high strength concrete; stress block; strength reduction factor; Eurocode; stress-strain

1. Introduction

The availability and advancement of material

technology and the acceptance has led to the

production of higher grades of concrete. High

strength concrete offers superior engineering

properties i.e. compressive strength, tensile strength,

durability, modulus of elasticity and overall better

performance when compared to the conventional

concrete [1,2]. However, high-strength concrete is

more brittle in nature because cracks in this material

do not always follow the aggregate-hardened cement

paste interfaces due to the improved interfacial bond

strength of high-strength concrete but may cut right

through the hardened cement paste and even the

aggregate particles leading to rapid propagation of

the cracks and sudden or sometimes explosive failure

of the concrete. Because of this problem, many

structural engineers hesitate in using high-strength

concrete, despite its obvious advantages. Research

on the behavior of HSC beams with concrete

strength higher than 55 MPa has been carried out in

the past and is still continuing, to understand the

behavior of HSC beams in flexure. Whilst there are

many publications proposing stress block models for

HSC beams, a universally accepted stress block

model is yet to be developed. In most design

standards, the conventional rectangular stress block

developed for Normal Strength Concrete is still

being used for design of HSC beams. Rectangular

stress block is generally used to calculate the

ultimate moment capacity of reinforced concrete

beams. The stress-strain curves for high strength

concrete are more linear than parabolic and hence it

was reasonable to infer that the rectangular stress

block parameters could be different.

The idea of using the equivalent rectangular

stress distribution was first proposed by Emperger [3]

and then modified by Whitney [4] for application to

ultimate strength design and later experimentally

verified by Hognestad et al. [5] and Mattock et al. [6].

To obtain accurate as well as well-controlled data on

flexure compression-loaded members, a test

procedure for a series of experiments on C-shaped

Brijesh Singh is a Manager at the Centre for Construction De-

velopment and Research National Council for Cement and

Building Materials

Corresponding author Vikas Patel is a Project Engineer at the

Centre for Construction Development and Research National

Council for Cement and Building Materials

P N Ojha is a Joint director at National Council for Cement

and Building Materials (NCB).

VV Arora is a joint director at the Centre for Construction De-

velopment and Research National Council for Cement and

Building Materials.

Journal of Asian Concrete Federation, Vol. 6, No. 1, June 2020

2

concrete specimens subjected to axial load and

bending moment was proposed by Hognestad et al.

and later was used by several researchers. The

position of neutral axis depth was kept fixed by

continuously monitoring strains on one surface of the

C-shaped specimen and adjusting the eccentricity of

the applied force so that the strains on the neutral

surface remain zero.

The rectangular stress block model was first in-

troduced by Hognestad et al. (1955) from experi-

mental work involving normal strength concrete.

Ashour [7] has shown that the flexural rigidity in-

creases as concrete compressive strength increases.

From the experimental study by Oztekin et al. [8], it

was observed that the rectangular stress block pa-

rameters used in ordinary concrete members cannot

be used safely for high strength concrete members.

Attard and Stewart (1998) [9] examined the applica-

bility of ACI 318-95 rectangular stress block param-

eters to high-strength concretes. They have shown

that for a ductile singly-reinforced rectangular sec-

tion, the ultimate moment capacity is relatively in-

sensitive to the stress block model. An experimental

study on the evolution of depth of neutral axis at fail-

ure with the ductility at bending on HSC beams was

carried out by Bernardo & Lopes (2004) [10].

The equivalent rectangular concrete stress block

adopted by various current RC design codes [11,12]

(European Committee for Standardization 2004;

Standards New Zealand 2006; ACI Committee 318-

2008) are depending only on the concrete strength.

However, from the comparison conducted by the Ho

et al. [13,19] using previous experimental test results

done by other researchers, the theoretical flexural

strengths predicted by RC design codes are signifi-

cantly smaller than the actually tested flexural

strengths. And from the results obtained by previous

researchers [14-17], it was found that the stress block

parameters were fairly scattered even though the

concrete strength is the same. Therefore, the assump-

tion of stress block parameters should depend on

other factors apart from concrete strength only.

It was found that the theoretical formulations

based on the use of the rectangular block diagram for

the concrete to compute the depth of neutral axis at

failure gave substantially smaller values as com-

pared to the experimental values. As such, it was

concluded that the rectangular stress block diagram

proposed by ACI 318-1989 was not adequate for

HSC beams. Cetin and Carrasquillo (1998) [18] re-

ported that no single equation of various codes and

research done in past seems to represent the flexural

strength of HSC with sufficient accuracy and, there-

fore, measured values should be used instead of

predicated ones.

Ultimate concrete compressive strength is an-

other important variable in the ultimate strength de-

sign. Although the ultimate flexural strength of rein-

forced concrete sections does not depend on this var-

iable, it can noticeably affect the ultimate curvature

of reinforced cross sections. Mattock et al. [6] con-

cluded that the value of 0.003 is a reasonably con-

servative value for ultimate strain of concrete. This

value has been accepted by many design codes (NZS

3101 2006; ACI 318-08 2008; AS 3600 2009). Kahn

et al. [20] reported that the ultimate value of 0.003 is

valid for concrete up to 102MPa and provided the

best prediction of the ultimate moment. According to

Mansur et al. study (1997), the maximum of 0.003

for concrete in compression may be extended to high

strength concrete. Ibrahim and MacGregor (1996)

results for ultimate concrete strain were considerably

higher than the limiting value of 0.003. However,

they concluded that based on the reported values in

previous tests of C-shaped specimens, the value of

0.003 used by the ACI code, seems appropriate as a

conservative lower bound of experimental data.

The paper focuses on calculation of stress block

parameters K (strength reduction factor) and k2 (fac-

tor for depth of resultant compressive force) from ex-

perimental strain values. In this study, the model pro-

posed in European design standard EC: 02-2004

have been analyzed and stress block parameters are

transformed to the parameters used in IS code design

procedure. Thereafter, comparison with the design

parameters are done with experimental results.

2. Concrete Ingredients

Crushed aggregate with a maximum nominal

size of 20 mm was used as coarse aggregate and nat-

ural riverbed sand confirming to Zone II as per IS:

383 was used as fine aggregate. Their physical prop-

erties are given in Table 1. The petrographic studies

conducted on coarse aggregate indicated that the ag-

gregate sample is medium grained with a crystalline

texture and partially weathered sample of granite.

The major mineral constituents were quartz, biotite,

plagioclase-feldspar and orthoclase-feldspar. Acces-

sory minerals are calcite, muscovite, tourmaline and

iron oxide. The petrographic studies of fine aggre-

gate indicated that the minerals present in order of

abundance are quartz, orthoclase-feldspar, horn-

blende, biotite, muscovite, microcline-feldspar, gar-

net, plagioclase-feldspar, tourmaline, calcite and

iron oxide. For both the coarse aggregate and fine

aggregate sample the strained quartz percentage and

their Undulatory Extinction Angle (UEA) are within

permissible limits. Feldspar grains are partially frac-

tured and shattered. The quality of both coarse and

fine aggregate is fair. The silt content in fine aggre-

gate as per wet sieving method is 0.70 percent.


3

Table 1 Properties of aggregates

Property Granite Fine Aggregate

20 mm 10 mm

Specific gravity 2.83 2.83 2.64

Water absorption (%) 0.3 0.3 0.8

Sieve

Analysis

Cumulative per-

centage

passing (%)

20 mm 98 100 100

10 mm 1 68 100

4.75 mm 0 2 95

2.36 mm 0 0 87

1.18 mm 0 0 68

600 µ 0 0 38

300 µ 0 0 10

150 µ 0 0 2

Pan 0 0 0

Abrasion, Impact & Crushing Value 19, 13, 19 - -

Flakiness % & Elongation % 29, 25 - -

One brand of Ordinary Portland cement (OPC

53 Grade) with fly ash and silica fume are used in

this study. The chemical and physical compositions

of cement OPC 53 Grade, Properties of fly ash and

silica fume are given in Table 2. Polycarboxylic

group based superplasticizer for w/c ratio 0.20, 0.27,

0.30 and 0.36 and Naphthalene based for w/c ratio

0.47 complying with requirements of Indian Stand-

ard: 9103 is used throughout the investigation. Water

complying with requirements of IS: 456-2000 for

construction purpose was used. The 3 days, 7 days

and 28 days compressive strength of cement OPC 53

Grade were 36.00 N/mm2, 45.50 N/mm2 and 57.50

N/mm2 respectively. The 28 days compressive

strength of controlled sample and sample cast with

flyash was 38.53 N/mm2 and 31.64 N/mm2 respec-

tively, when testing was done in accordance with IS:

1727. The 07 days compressive strength of con-

trolled sample and sample cast with silica fume was

12.76N/mm2 and 14.46 N/mm2 respectively, when

testing was done in accordance with IS: 1727.

3. Mix Design Details

In this study, the four different mixes ranging

from w/c ratio 0.47 to 0.20 using granite aggregate

were studied for determining short term mechanical

properties of High Strength Concrete. For each type

of aggregate, three separate batches were prepared.

Table 2 Physical, chemical and strength characteristics of cement

Characteristics OPC -53 Grade Silica Fume Fly Ash

Physical Tests:

Fineness (m2/kg) 320.00 22000 403

Soundness Autoclave (%) 00.05 - -

Soundness Le Chatelier (mm) 1.00 - -

Setting Time Initial (min.) &

(max.)

170.00 & 220.00 - -

Specific gravity 3.16 2.24 2.2

Chemical Tests:

Loss of Ignition (LOI) (%) 1.50 1.16 -

Silica (SiO2) (%) 20.38 95.02 -

Iron Oxide (Fe2O3) (%) 3.96 0.80 -

Aluminium Oxide (Al2O3) (%) 4.95 - -

Calcium Oxide (CaO) (%) 60.73 - -

Magnesium Oxide (MgO) (%) 4.78 - -

Sulphate (SO3) (%) 2.07 - -

Alkalies (%) Na2O & K2O 0.57 & 0.59 -

Chloride (Cl) (%) 0.04 - -

IR (%) 1.20 - -

Moisture (%) - 0.43 -


4

The slump of the fresh concrete was kept in the range

of 75 - 100 mm. A pre-study was carried out to de-

termine the optimum superplasticizer dosage for

achieving the desired workability based on the slump

cone test as per Indian Standard. The mix design de-

tails of specimens are given in Table 3. Adjustment

was made in mixing water as a correction for aggre-

gate water absorption. For conducting studies, the

concrete mixes were prepared in pan type concrete

mixer. Before use, the moulds were properly painted

with mineral oil, casting was done in three different

layers and each layer was compacted on vibration ta-

ble to minimize air bubbles and voids. After 24

hours, the specimens were demoulded from their re-

spective moulds. The laboratory conditions of tem-

perature and relative humidity were monitored dur-

ing the different ages at 27±2°C and relative humid-

ity 65% or more. The specimens were taken out from

the tank and allowed for surface drying and then

tested in saturated surface dried condition.

4. Stress Strain Study on High Strength

Concrete and Normal Strength Concrete

For stress strain characteristics of the high

strength concrete, concrete specimens were tested in

a closed-loop servo hydraulic compression testing

machine of 3000 kN capacity. Two extensometers at

the middle half of the height were used to get strain

and two strains were averaged. To obtain a full

stress-strain curve, a slow rate of loading in the range

of 1300 to 1500 N/sec was adopted for a whole com-

pression test. In general, the normal strength con-

crete gradually fails after reaching its peak load, but

the high strength concrete suddenly explodes at peak

load. Strain at peak stress and ultimate strain were

recorded for further analysis (Table 4).

4.1 Determination of stress block parameters

from experimentally obtained strain values

The total compressive force Cu and its location

below the top fibre can be expressed in terms of

stress block factors k1, k2 and k3.

k1= shape factor = ratio of the area of stress block

ABCD to area of rectangle AFCD

k2= ratio of depth of resultant compressive force to

depth of neutral axis(X)

r1 = 𝐴𝐸

𝐴𝐷=

𝜀𝑐

𝜀𝑐𝑢 and r2 =

𝐸𝐷

𝐴𝐷=

𝜀𝑐𝑢−𝜀𝑐

𝜀𝑐𝑢,

𝜀𝑐 = strain after which concrete yields at constant

stress of (αcc×S1) fck

𝜀𝑐𝑢 = ultimate strain in concrete

αcc = factor for consideration of long term effects

including the way load is applied = 0.85

S1 = factor for conversion of cube to cylinder

strength

Here, k1 is calculated by calculating the area of the

shaded portion and dividing by the area of the rec-

tangle, thus represents the ratio of the area of stress

Table 3 Concrete mix design details for study done

w/c

Total Cementitious Content

[Cement C + Fly ash (FA) +

Silica Fume (SF)] (Kg/m3)

Water

Content

(kg/m3)

Admixture

% by

weight of

cement

Fine aggregate as

% of total aggre-

gate by weight

28-Days

strength of con-

crete (N/mm2)

0.47 362 (290+72+0) 170 1.00 35 45.72

0.36 417 (334+83+0) 150 0.45 39 68.57

0.27 525 (400+75+50) 140 0.70 39 88.60

0.20 750 (548+112+90) 150 1.75 35 97.76

Table 4 Strain at peak stress and ultimate strain recorded

Cyd str,

N/mm2 Cube/Cyd

Cube Str,

N/mm2

Ec,

micro-strain

Ecu,

micro-Strain

24.00 1.28 30.72 2284 3711

23.50 1.28 30.08 1972 3789

34.40 1.28 44.03 2175 3063

33.80 1.28 43.26 1877 3220

48.60 1.24 60.26 2151 3324

46.09 1.24 57.15 2341 3323

76.83 1.16 89.12 2702 2931

76.18 1.16 88.37 2539 2729

103.90 1.14 118.45 2774 2774

106.00 1.14 120.84 2799 2799


5

For Concrete up to M55 For Concrete above M55 to M90

For calculation of area ABE parabola is

considered

For calculation of area ABE, Triangle is considered as for

grade between M55 and M90 in actual stress diagram the

shape of area ABE is somewhere between linear and parabola

k1= 2

3 r1 + r2

k2 = [2

3 r1 (r2 +

3

8 r1) + r2 (

r2

2)]/k1

k3 = αcc×S1

k1= 1

2 r1 + r2

k2 = [1

2 r1 (r2 +

1

3 r1) + r2 (

r2

2)] /k1

k3 = αcc×S1

block ABCD to the area of rectangle AFCD. k2 is

calculated by taking moment of shaded areas about

axis CD and equating it to the moment of resultant

force Cu about axis CD. K3 is the stress reduction

factor calculated by considering αcc = 0.85 and con-

version factor S1.

Cu = b × area ABCD = b × k1 × area AFCD

Cu = b × k1 × k3 × fck × X

Partial factor of safety, γmc = 1.5 for concrete & γs =

1.15 for steel

Cu = k1 × b× Xu × k3∗fck

γmc

Cu = K× fck × b × Xu (1)

where, K = k1 ∗k3

γmc (2)

Xu

𝑑=

fyγs

∗ Ast

𝐾 ∗ fck ∗ b ∗ d

Mu = Cu × (d-k2 × X) (Compression) (3)

Mu = Tu × (d-k2 × X) where Tu = fy

γs × Ast (Ten-

sion) (4)

The two basic factors K and k2 for flexural de-

sign are worked out from the strain values recorded

experimentally for different strength of concrete

based on the above method (Table 5).

Transformation of equation of Euro-code into IS

code format (basic flexural design factors)

Euro-code uses different philosophy for deter-

mination of compressive force in a section and there-

fore has different factors. These factors were clubbed

together to form the representative equation similar

to IS code equation for calculation of compressive

force to compare the reduction factor K and factor to

calculate lever arm k2.

Cu = λ × ƞ × fcd × b × Xu

Table 5 Calculation of K and k2 as per the IS code approach with experimentally obtained strain values

Cyd str Cube/Cyd Cube Str Ec Ecu r1 r2 k1 S1 k3 Ymc K(IS cur) k2

24.00 1.28 30.72 2284 3711 0.62 0.38 0.79 0.78 0.66 1.5 0.35 0.41

23.50 1.28 30.08 1972 3789 0.52 0.48 0.83 0.78 0.66 1.5 0.37 0.42

34.40 1.28 44.03 2175 3063 0.71 0.29 0.76 0.78 0.66 1.5 0.34 0.40

33.80 1.28 43.26 1877 3220 0.58 0.42 0.81 0.78 0.66 1.5 0.36 0.41

48.60 1.24 60.26 2151 3324 0.65 0.35 0.68 0.81 0.69 1.5 0.31 0.36

46.09 1.24 57.15 2341 3323 0.70 0.30 0.65 0.81 0.69 1.5 0.30 0.36

76.83 1.16 89.12 2702 2931 0.92 0.08 0.54 0.86 0.73 1.5 0.26 0.34

76.18 1.16 88.37 2539 2729 0.93 0.07 0.53 0.86 0.73 1.5 0.26 0.33

103.90 1.14 118.45 2774 2774 1.00 0.00 0.50 0.88 0.75 1.5 0.25 0.33

106.00 1.14 120.84 2799 2799 1.00 0.00 0.50 0.88 0.75 1.5 0.25 0.33


6

Euro-code equation for compressive force

fcd = {αcc × S1 / γcc} × fck

Cu = λ × ƞ × {αcc × S1 / γcc} × fck × b × Xu (5)

Compare equation (i) and (iii) and assume, K’ = λ ×

ƞ × αcc × S1 / γcc (6)

Cu = K’ × fck × b × Xu

Euro-code equation in IS code format

k2 = λ/2

The factor K and k2 are directly related to the

bandwidth between strain at peak stress and ultimate

strain of the concrete as well as the shape of the stress

strain curve of the concrete. The strength reduction

factor K reduces as the bandwidth between strain at

peak stress and ultimate strain decreases with in-

crease in strength and also the stress strain curve be-

comes steeper or linear for higher grades of concrete.

4.2 Calculation of moment capacity for balanced

section and comparison with that obtained as per

Euro-Code

To understand the effect of ultimate strain val-

ues on the moment capacities of the flexural mem-

bers moment capacities were calculated for a bal-

anced section from stress block parameters derived

from experimentally obtained strain values and were

also compared with the moment capacities calcu-

lated as per Euro-code. A representative section with

dimension (b=200 mm and D=400 mm with a clear

cover of 25 mm) was used for calculation of moment

capacity.

IS method:

Xu(max)

𝑑=

∈𝑐

∈𝑐+ ∈𝑠𝑢

where, Ɛc is ultimate strain in top most compression

fiber; Ɛsu is ultimate strain in steel = 0.002 +

0.87fy/Es

Mu (KN-M) = Cu × (d - k2 × X)

Mu(max) = K × fck × b × Xu(max) × (d - k2 × Xu(max)) for

balanced section

Euro-code method:

K4 = 1.25 × (0.6 + (0.0014 × 1000000 / Ɛcu))

Xu(max)

𝑑 = ((1 - 0.44) / k4) for cylindrical strength less

than 50 MPa

Xu(max)

𝑑 = ((1 - 0.54) / k4) for cylindrical strength

more than 50 MPa

Mu = Cu × (d - (λ/2) × X)

Mu(max) = K’ × fck × b × Xu × Xu(max) × (d - (λ/2) ×

Xu(max)) for balanced section

The ultimate strain of concrete has direct impact

on depth of neutral axis for balanced section which

is directly related to the maximum capacity of the

member. The ultimate strain values decreased as the

strength of concrete increases and becomes nearly

constant after 90 N/mm2 as seen from experimental

values and also as per many international standards.

Current IS code gives a constant value of ultimate

strain of 0.0035 up to concrete grade M50. This

Table 6 Calculation of K and k2 as per the Euro code transformed in IS code parameters

Cyd str

N/mm2 Cube/Cyd

Cube Str,

N/mm2

Ec,

micro-

strain

Ecu,

micro-

strain

λ ƞ αcc S1 Ycc K'(EC) k2

24.00 1.28 30.72 2000.00 3500.00 0.80 1.00 0.85 0.78 1.5 0.35 0.40

23.50 1.28 30.08 2000.00 3500.00 0.80 1.00 0.85 0.78 1.5 0.35 0.40

34.40 1.28 44.03 2000.00 3500.00 0.80 1.00 0.85 0.78 1.5 0.35 0.40

33.80 1.28 43.26 2000.00 3500.00 0.80 1.00 0.85 0.78 1.5 0.35 0.40

48.60 1.24 60.26 2000.00 3500.00 0.80 1.01 0.85 0.81 1.5 0.37 0.40

46.09 1.24 57.15 2000.00 3500.00 0.81 1.02 0.85 0.81 1.5 0.38 0.40

76.83 1.16 89.12 2485.95 2610.53 0.73 0.87 0.85 0.86 1.5 0.31 0.37

76.18 1.16 88.37 2479.67 2612.77 0.73 0.87 0.85 0.86 1.5 0.31 0.37

103.90 1.14 118.45 2600.00 2600.00 0.67 0.73 0.85 0.88 1.5 0.24 0.33

106.00 1.14 120.84 2600.00 2600.00 0.66 0.72 0.85 0.88 1.5 0.24 0.33


7

value of 0.0035 is valid upto M35 grade of concrete

only as per the experimental results. Based on trend

shown by experimental results the ultimate strain

value given in Euro Code EC-02 seems to be more

realistic for higher grade concrete whereas constant

value of ultimate strain for concrete grade between

M25 to M50 (cylindrical strength) does not seems to

be realistic. Therefore, the value of ultimate strain in

concrete should be restricted to a lower value for

higher grades of concrete.

5. Conclusions

The method or approach worked out in the pa-

per will be useful in understanding the basic flexural

design philosophy. The stress-strain characteristics

or the effect of stress-strain parameters on the flex-

ural design equations are the fundamental of stress

block parameters and same needs to be derived from

the strain at peak stress and ultimate strain. The fac-

tor K (strength reduction factor) and k2 (factor for

the depth of resultant compressive force) are directly

related to the bandwidth between strain at peak stress

and ultimate strain of the concrete as well as the

shape of the stress-strain curve of the concrete. The

strength reduction factor K reduces as the bandwidth

between strain at peak stress and ultimate strain de-

creases with an increase in strength and also the

stress-strain curve becomes steeper or linear for

higher grades of concrete. The ultimate strain of con-

crete has a direct impact on the depth of the neutral

axis for a balanced section which is directly related

to the maximum capacity of the member.

The ultimate strain values decrease as the strength of

concrete increases and become nearly constant after

90 N/mm2 as seen from experimental values and also

Table 7 Calculation of moment capacity as per the IS code approach with experimentally obtained strain val-

ues for balanced section

Cyd str.

N/mm2 Cube/Cyd

Cube

Str

N/mm2

Ɛc,

micro-

strain

Ɛcu,

micro-

strain

Ɛsu,

micro-

strain

K

(IS cur) k2 Xu(max)/d

Mo-

ment

(kN-M)

24.00 1.28 30.72 2284.00 3711.00 4175 0.35 0.41 0.471 109.35

23.50 1.28 30.08 1972.00 3789.00 4175 0.37 0.42 0.476 111.49

34.40 1.28 44.03 2175.00 3063.00 4175 0.34 0.40 0.423 139.38

33.80 1.28 43.26 1877.00 3220.00 4175 0.36 0.41 0.435 146.72

48.60 1.24 60.26 2151.00 3324.00 4175 0.31 0.36 0.443 184.52

46.09 1.24 57.15 2341.00 3323.00 4175 0.30 0.36 0.443 168.28

76.83 1.16 89.12 2702.00 2931.00 4175 0.26 0.34 0.412 222.27

76.18 1.16 88.37 2539.00 2729.00 4175 0.26 0.33 0.395 210.97

103.90 1.14 118.45 2774.00 2774.00 4175 0.25 0.33 0.399 271.46

106.00 1.14 120.84 2799.00 2799.00 4175 0.25 0.33 0.401 278.21

103.90 1.14 118.45 2600.00 2600.00 4175 0.25 0.33 0.384 262.51

106.00 1.14 120.84 2600.00 2600.00 4175 0.25 0.33 0.384 267.82

Table 8 Calculation of moment capacity as per the Euro code (transformed in IS code parameters) for bal-

anced section

Cyd str.

N/mm2 Cube/Cyd

Cube

Str,

N/mm2

Ɛc,

micro-

strain

Ɛcu,

micro-

strain

k4 K'(EC) λ/2 Xu(max)/d

Mo-

ment

(kN-M)

24.00 1.28 30.72 2000.00 3500.00 1.25 0.35 0.40 0.448 106.60

23.50 1.28 30.08 2000.00 3500.00 1.25 0.35 0.40 0.448 104.38

34.40 1.28 44.03 2000.00 3500.00 1.25 0.35 0.40 0.448 152.79

33.80 1.28 43.26 2000.00 3500.00 1.25 0.35 0.40 0.448 150.13

48.60 1.24 60.26 2000.00 3500.00 1.25 0.37 0.40 0.368 186.19

46.09 1.24 57.15 2000.00 3500.00 1.25 0.38 0.40 0.368 179.93

76.83 1.16 89.12 2485.95 2610.53 1.42 0.31 0.37 0.324 210.12

76.18 1.16 88.37 2479.67 2612.77 1.42 0.31 0.37 0.324 209.60

103.90 1.14 118.45 2600.00 2600.00 1.42 0.28 0.35 0.323 251.85

106.00 1.14 120.84 2600.00 2600.00 1.42 0.28 0.35 0.323 256.94


8

as per many international standards.

The study indicates that it is better to give ide-

alized stress strain curve for different concrete

grades in Design Standards which will also highlight

the decrease in the bandwidth of strain values with

increase in the grade of concrete and will take into

effect the steepness of curve also. Therefore, it is im-

portant to include the influence of stress-strain char-

acteristics i.e., the shape of the stress-strain curve,

the strain at peak stress and ultimate strain in the

flexural design for safe and efficient design of struc-

tural members using high strength concrete.

References

1. V.V. Arora; Brijesh Singh; and Shubham Jain

(2017) “Effect of indigenous aggregate type on

mechanical properties of High Strength

Concrete”, Indian Concrete Journal, 91(1):34-

44.

2. V.V. Arora; Brijesh Singh; and Shubham Jain

(2016) “Experimental studies on short term

mechanical properties of High Strength

Concrete”, Indian Concrete Journal, 90(10):65-

75

3. Emperger, F. (1904) "Ein graphischer Nachweis

der Tragfahigkeit und aller in einem Tragwerke

aus Eisenbeton auftretenden Spannungen",

Beton und Eisen 4(5):306-320

4. Whitney, C. S. (1937) “Design of reinforced

concrete members under flexure or combined

flexure and direct compression”, ACI Journal

Proceedings, ACI

5. Hognestad, E., N. W. Hanson, et al. (1955)

“Concrete stress distribution in ultimate strength

design”, ACI Journal Proceedings, ACI

6. Mattock, A. H., L. B. Kriz, et al. (1961)

“Rectangular concrete stress distribution in

ultimate strength design”, ACI Journal

Proceedings, ACI

7. Ashour, S. A. (2000) “Effect of compressive

strength and tensile reinforcement ratio on

flexural behavior of High-Strength Concrete

Beams”, Engineering Structures, Vol.22, p.413-

423

8. Oztekin E.; Pul S.; Husem, M. (2003)

“Determination of rectangular stress block

parameters for High Performance Concrete”,

Engineering Structures, 25:371-376

9. Attard M.M.; Stewart M.G. (1998) “Two

parameter stress block for High-Strength

Concrete”, ACI Structural Journal 95(3):305–

317

10. Bernardo, L.F.A.; Lopes S.M.R. (2004)

“Neutral axis depth versus flexural ductility in

High-Strength Concrete Beams”, J. Struct. Eng.,

130(3):452-459

11. ACI Committee 318 (2008) “Building Code

Requirements for Reinforced Concrete and

Commentary ACI 318M-08”, Manual of

Concrete Practice. American Concrete Institute.

Michigan. pp 465. USA

12. Standards New Zealand, NZS 3101 (2006),

“Concrete Structures Standard, Part 1 - The

Design of Concrete Structures”, Wellington,

New Zealand

13. Ho JCM; Pam HJ; Peng J; Wong YL (2010)

“Maximum concrete stress developed in

flexural RC members”, Computers and

Concrete

14. Swartz SE; Nikaeen A; Narayan BHD;

Periyakaruppan N; Refai TME (1985)

“Structural bending properties of high strength

concrete”, ACI Special Publication, 87(9):

p.147-178

15. Tan TH; Nguyen NB (2005) “Flexural behavior

of confined high-strength concrete columns”,

ACI Structural Journal. 102(2):198-205

16. Mansur MA; Chin MS; Wee TH (1997)

“Flexural behavior of high-strength concrete

beams”, ACI Structural Journal. 94(6):663-673

17. Kaar PH; Hanson NW; Capell HT (1978)

“Stress-strain characteristics of high strength

concrete”, ACI Special Publication. Douglas

McHenry International Symposium on Concrete

and Concrete Structures. ACI. Detroit. pp.161-

185

18. Cetin, A.; Carrasquillo, R. L. (1998) “High-

Performance Concrete: Influence of Coarse

Aggregates on Mechanical Properties”, Journal

of ACI Materials, 95:252-261

19. Peng Ho; Pam Wong (2012) “Equivalent stress

block for normal-strength concrete

incorporating strain gradient effect”, Magazine

of Concrete Research Volume 64 Issue

20. Kahn, L. F.; K. F. Meyer (1995) "Rectangular


9

stress block for nonrectangular compression

zone", ACI Structural Journal 92(3)

21. Ibrahim, H. H. H.; J. G. MacGregor (1996)

"Flexural behavior of laterally reinforced high-

strength concrete sections", ACI Structural

Journal 93(6)

22. Eurocode 2-2004: Design of concrete structures,

European Committee for Standardization

23. IS: 456-2000, “Code of Practice for Plain and

Reinforced Concrete”, Bureau of Indian

Standards, New Delhi

24. IS: 383-2016, “Coarse and fine aggregate for

Concrete-Specification”, Bureau of Indian


25. CPWD-2009, Central Public Works Department

Manual, Government of India

26. IS: 1727-1967, “Methods of Test for Pozzolanic

Materials”, Bureau of Indian Standards, New

Delhi

27. IS: 516-2018 "Hardened concrete-method of

tests", Bureau of Indian Standards, New Delhi


Vol. 6, No. 1, pp. 10-25, June 2020

ISSN 2465-7964 / eISSN 2465-7972

https://doi.org/10.18702/acf.2020.6.6.10

10

Technical Paper

Effect of clay as deleterious material on properties of nor-

mal-strength concrete

Harish R. Choudhary, Saha Dauji*, Arham Siddiqui

(Received February 12, 2020; Revised July 2, 2020; Accepted July 3, 2020; Published July 3, 2020)

Abstract: Sustainability concerns prompted use of crushed aggregates in concrete, wherein deleterious

materials might get inadvertently included. Some deleterious materials are allowed up to limiting values

by most standards, which, however, are silent about the quantification of their effects on properties of

concrete – which would be site specific. For an important Indian infrastructure, this study quantifies

effects of clay fines as deleterious material in concrete, on workability (slump) and strength (cube com-

pression; split tensile; flexural tensile tests) around the limit (1% of fine aggregates by weight) stipulated

by the Indian standard. The novelty of the study is that, contrary to the literature in this domain, the

effects are studied without alteration of the mix proportions – a different practical scenario. The limit of

clay fines in concrete allowed by Indian standard was found to be adequate considering strength param-

eters, but for maintaining target workability, the limit would be revised to 0.5% of the fine aggregates.

Generally, the variations of concrete properties with the increasing clay fines were: (1) the workability

and split tensile strength reduced monotonically, in non-linear fashion; (2) compressive strength (beyond

7 days) and the flexural tensile strength (modulus of rupture) reduced monotonically in linear manner.

Keywords: concrete; clay; workability; compressive strength; modulus of rupture; split tensile strength

1. Introduction

During last four to five decades, concrete has

grown in popularity as a building material

particularly due to its flexibility in the geometry, its

resistance to the environment and fire as compared

to steel, and cost effectiveness. The concrete

generally consists of the ingredients: coarse

aggregate, fine aggregate, cement and other binding

materials, water, and admixtures, as applicable. The

design of the concrete mix is performed with the

national building codes [1,2] in order to achieve the

desired characteristic strength. The mix design,

being conducted in the laboratory, happens under

strict quality control. However, the actual concrete

being mixed either at site (more uncertainty) or at the

batching plant (less uncertainty and better control

than site mix), there always exist the possibilities of

deleterious materials being inadvertently mixed in

concrete. In fact, Indian and International codes [3,4]

permit certain limiting amount of deleterious

material being present in concrete, and still it can be

acceptable. However, the standards are silent on the

effects of the deleterious material, if present, on the

properties of green or hardened concrete.

The deleterious material is defined in literature

as materials which might affect the concrete in the

following ways:

1) Materials which might interfere with the

process of hydration of cement [5];

2) Coatings which would prevent development

of good bond between aggregate and cement paste

[5];

3) Weak or unsound materials which would

affect the strength of concrete [5].

The classification of deleterious materials was

conducted as early as 1950-s [6]. Broadly classified,

the deleterious materials would fall under the follow-

ing categories:

1) Organic impurities: they interfere with the

hydration process of cement. They are generally pre-

sent in the form of humus or organic loam and mostly

found in the fine aggregates [5,6].

2) Clay and other fine materials: Clay is usually

present in the form of coating on the aggregates and

interferes with bond between aggregates and cement

matrix. Silt and other fine materials might be present

as coatings, when the bond is affected; or as loose

Harish R. Choudhary is a Scientific Officer at Nuclear Fuel

Complex, Kota, India.

Corresponding author Saha Dauji is a Scientific Officer at

Bhabha Atomic Research Centre, Mumbai, India and a lecturer

at Homi Bhabha National Institute, Mumbai, India

Arham Siddiqui is a is a Scientific Officer at Nuclear Fuel

Complex, Kota, India.


11

particles which are not bonded and reduces the

strength and durability of concrete [5,6].

3) Salt contamination: mostly occurs when the

fine aggregate (sand) originates from sea and is a ma-

jor durability concern as it could effect chloride in-

duced corrosion of reinforcement [5,6].

4) Unsound particles: these are particles which

might lose their integrity after casting of concrete or

might undergo disruptive expansion in contact with

water or in case of freezing. Clay lumps, if present in

concrete, would be considered as the unsound parti-

cles [5,6].

In developing countries, the issue of utilization

of locally available aggregates for concrete has been

examined [7]. The effect of presence of clay in con-

crete has been investigated by many researchers

across the world [8-21], and techniques have been

proposed to evaluate the possible harmful influences

of the clay in concrete [21].

There has been a growing concern of the eco-

logical and environmental impacts of the large scale

quarrying or dredging conducted for the supply of

aggregate as raw materials for concrete. These in-

clude the destruction of habitat, disruption of food

chains, disturbance to the local ecology, instability

of the courses of rivers, increased erosion, scour, or

depositions on the banks of rivers or the coasts due

to sand dredging, and changes in the river regimes or

coastal hydrodynamics [22]. Therefore, researchers

all over the world are exploring alternate sources of

aggregate such as recycled concrete, building mate-

rials, glass, tyres or ceramic tiles [22-25]. The dis-

posal of the excavated materials from major con-

struction projects pose another concern which often

becomes extremely expensive because of the large

lead distances involved between the construction

sites and the disposal site. As a practical solution to

either of the aforementioned problems, in many pro-

jects, the practice of using crushed aggregates from

the excavated rock from the same site is presently

getting preference from sustainability considerations

[26-28].

Till recently, use of crushed fine aggregates was

less popular in Indian construction projects, particu-

larly due to the lack of experience and data. But ex-

perience from the earlier works and stipulations of

the national codes [3] have fostered use of up to

100% crushed coarse and fine aggregate manufac-

tured from the locally excavated rock, provided they

met the quality standards [3]. In a major infrastruc-

ture project in India, the same scheme was adopted

and the excavated rock (Fig. 1a) from the site was

crushed in the local crusher plants (Fig. 1b) for sub-

sequent use as aggregates, both coarse and fine, in

normal-strength concrete. The geology of the site

showed that the major rock type was sandstone with

clay matrix. This could lead to contamination of

crushed aggregates with clay. Clay being an unsound

particle would act like weak intrusion in concrete

mix. The standard practice of preparation of aggre-

gates includes washing procedure, precisely for re-

moval of such impurities. However, the possibility

of clay in the concrete made with the crushed aggre-

gates from the excavated rock could not be ruled out.

It has been mentioned earlier that the national code

allows a certain amount of clay in the concrete (1%

for India [3]), but does not specify the effect on the

workability or strength of concrete. There are vari-

ous limits of clay in concrete based on its mineral

composition available in literature [20]. This is un-

derstandable as the deleterious materials, in this case

– clay, would vary in properties from site to site and

consequently, the effect on concrete properties could

also be different. The deleterious materials being site

specific, the study on the effect of the same on con-

crete would also be site specific. As the project was

of national importance, the possible effects of inad-

vertent inclusion of clay on the properties of fresh

and hardened concrete were deemed necessary to be

critically examined.

(a) (b)

Fig 1: (a) excavated rock at site; (b) crusher plant at site


12

Schuster [8] presented a detailed review of the

studies pertaining to the effects of various deleteri-

ous materials present in concrete aggregates on the

properties of concrete. The effect of very fine aggre-

gate was reported as increase in the 91-day strength,

at constant workability [9]. Though some researchers

[29] indicated that very large amount of fines could

be incorporated into concrete without much detri-

mental effect on its properties. Abou-Zeid and

Fakhry [10] reported that addition of clay fines to

concrete in even small amount reduced the workabil-

ity to a significant extent, though small proportions

of clay increased the compressive strength by a cer-

tain extent. Gullerd [11] examined the effects of var-

ious aggregate coatings on the performance of Port-

land cement concrete, including durability (drying

shrinkage, freeze-thaw, chloride vulnerability) and

strength (compressive) and concluded that the clay

coatings were more detrimental compared to the dust

or carbonates. Munoz et al. [12] conducted detailed

experimentation on the effect of clay coatings on ag-

gregates and reported that fraction of clay getting de-

tached and entering water phase would depend on

the original clay content as well as the properties of

clay, which could vary from site to site. The clay in

cement paste affected the rate of hydration, and this

effect would also be dependent on the nature of clay,

they indicated. Katz and Baum [13] explored the

higher levels of fines on the strength properties of

concrete with detailed experimental program, but us-

ing concrete of constant workability. They con-

cluded that as long as workability was controlled to

the target values by addition of suitable admixtures,

the strength of concrete increased up to 30% with in-

creases in volume change of fresh and hardened con-

crete – with addition of small amounts of fines.

Munoz et al. [14] examined the effects of vari-

ous aggregate coatings and films on the concrete per-

formance, and concluded that the clay coatings were

more harmful than other microfine mineralogy such

as dust or carbonates, the observations being similar

to earlier literature [11]. Cramer [30] indicated that

the stipulations of the codes might not be adequate in

assuring the desired properties of concrete when mi-

crofines were concerned, due to their typical miner-

alogical composition. They investigated the effect of

microfines on strength and workability of concrete,

and concluded that dolomitic microfines did not af-

fect workability much, but clay microfines definitely

affected the workability adversely with a little im-

provement in strength properties. Abib et al. [15] ex-

plored the properties of self-compacting concrete

with fine clay from waste crushed brick, and con-

cluded that the strength improved with up to 5% of

the fine clay addition.

The earlier studies examined the strength prop-

erties of concrete at constant workability – a condi-

tion which might not actually exist at a construction

site. Hence, in this study, the concrete mix is kept

unaltered, and the effect of clay on both workability

and the strength of concrete are studied for different

proportions of clay in concrete. To the knowledge of

the authors, this approach has not been earlier re-

ported in literature. The initial results of the experi-

ments were presented by the authors in a conference

earlier [16] and the analysis of the complete set of

results is presented in this article.

This article presents the investigation conducted

with slump test for the workability of fresh concrete

and a series of cube and cylinder tests for the strength

properties of hardened concrete at different ages

from 7 days to 91 days. The effect of clay as delete-

rious material in concrete was investigated around

the acceptable limit stipulated by the Indian code [3].

The results of this study would help to reaffirm the

limit stipulated in the code as acceptable for the site,

or would help to fix site specific limit on the clay in

aggregates for attaining the desired strength and

workability of the resulting concrete.

2. Methodology

2.1 Standard mix design

The grade of concrete for the facility was M25,

that is, the 95 percent confidence value of 28-day

cube compressive strength was 25 MPa. The target

slump value was 120 mm, measured according to the

Indian standard [31]. The design of the concrete mix

was carried out at the site according to the Indian

standard [2], and the final proportions for the M25

mix is reproduced in Table 1 for ready reference. The

target strength of the mix thus prepared would be

31.6 MPa.

Table 1 Standard design mix (M25)

Grade of

concrete W/C ratio

Quantities of materials per cubic meter of concrete

Slump

(mm) Cement

(kg)

Water

(L)

Plasticizer

(kg)

Sand

(kg)

Coarse aggregate

(kg)

20 mm 10 mm

M25 0.48 334 160 2 806 624 416 120


13

The crushed aggregates used in the mix design,

namely, 20 mm, 10 mm, and crushed sand are de-

picted in Fig. 2(a), Fig 2(b) and Fig. 2(c) respec-

tively.

2.2 Properties of deleterious material: clay

The origin of the clay at the site can be traced

back to the Vindhyan range of sedimentary layer of

reddish black soil [32]. The clay fines have a liquid

limit of 42% and plasticity index of 27%, which clas-

sifies the fines as clay of medium to high expan-

sion/compression properties, according to Indian

standard [33]. The clay fines have a free swell index

of 53% lying in the range of very high degree of ex-

pansiveness. Particularly, this property might cause

porosity in concrete and this might eventually lead to

strength and durability issues [11,12,14,30].

2.3 Design of experiment: concrete mix with clay

The IS code [3] stipulation is 1% as a maximum

for clay to be present in concrete, as a percentage of

the crushed fine aggregate by weight. The standard

mix (SM) contains no clay, and two mixes are taken

between standard mix (SM) and the IS limit. Two

other mixes are taken higher than the IS code stipu-

lation, namely, 1.5 times and 2 times the limiting

value. All the details of the mixes are reproduced in

the Table 2. The workability (slump) was tested for

all the mixes before the casting of the cubes and cyl-

inders. The strength tests were conducted at different

ages of concrete from 7 days to 91 days and the spe-

cific details are listed in Table 3. The cube strength

was tested at 7 days, 14 days, 28 days, 63 days and

91 days whereas the cylinder and flexural tests were

conducted at 28 days only.

The IS code [3] stipulation is 1% as a maximum

for clay to be present in concrete, as a percentage of

the crushed fine aggregate by weight. The standard

mix (SM) contains no clay, and two mixes are taken

between standard mix (SM) and the IS limit. Two

other mixes are taken higher than the IS code stipu-

lation, namely, 1.5 times and 2 times the limiting

value. All the details of the mixes are reproduced in

the Table 2. The workability (slump) was tested for

all the mixes before the casting of the cubes and cyl-

inders. The strength tests were conducted at different

ages of concrete from 7 days to 91 days and the spe-

cific details are listed in Table 3. The cube strength

was tested at 7 days, 14 days, 28 days, 63 days and

91 days whereas the cylinder and flexural tests were

conducted at 28 days only.

(a) (b)

(c) (d)

Fig 2: Aggregates (a) 20 mm; (b) 10 mm; (c) crushed sand; (d) deleterious material: prepared clay before be-

ing pulverized


14

2.4 Standard procedure for mixing, casting and

curing of concrete specimens

The experiments are targeted towards capturing

the effects of deleterious materials which comes to

around a thousandth by weight. Hence, it was essen-

tial to adopt a carefully formulated Standard Proce-

dure for conducting the experiments, which is briefly

discussed in the following sub-sections. These are

conforming to the stipulations laid down by the In-

dian standards [1,34,35].

2.4.1 Preparation of materials

All materials are brought to room temperature,

that is, within 27ºC ± 3º C, before commencing with

the mixing. The cement samples, on arrival at the la-

boratory, were thoroughly mixed dry by hand so as

to ensure the greatest possible blending and uni-

formity in the material. The cement was stored in a

dry place, in air-tight metal containers. Aggregates

for each batch of concrete were air dried before using

in mix. Water was potable water and was at room

temperature. The preparation of the clay, to be added

as deleterious material deserves particular mention.

The soil was sieved through 75 micron sieve in wet

condition (slurry) to collect the clay fines in the pan.

This clay slurry is oven-dried and then pulverized to

obtain the clay for mixing with concrete during the

preparation of various concrete mixes.

2.4.2 Proportioning of materials

The proportioning of the materials was done by

weight, per cubic meter of concrete according to the

mix design listed in the Table 2 for the particular

mix. Weigh batching was used for proportioning.

The weighing of the cement and each size of aggre-

gate is done to an accuracy of 0.1 percent of the total

weight of each batch.

2.4.3 Mixing

A small mixing machine was used for mixing of

materials. Hand loading was adopted for loading the

materials on the mixer. The mixing drum was loaded

with about one-half of the coarse aggregates, fol-

lowed by the fine aggregates, the deleterious mate-

rial: clay, the cement and finally with the remaining

coarse aggregate on top, in that order. The water, in

which the admixture was already mixed, was added

to the mixing drum just before starting the mixing

machinery. The period of mixing was kept not less

than 2 minutes and was continued till the resulting

concrete was uniform in appearance.

2.4.4 Casting of cubes, beams and cylindrical

specimens

The size of the cube test specimens was 150 mm

× 150 mm × 150 mm, confirming to Indian standard

[36]. The average strength of three cubes was re-

ported as the compressive strength at a particular age

for the mix, except for 28 days' strength – whence

the number of cubes was 13. For the six different

mixes (CM, and C1 to C5) and five different ages (7,

14, 28, 63 and 91 days) for test of compressive

strength, the total number of cube specimens tested

was 150 (Table 3). The cylindrical moulds were of

150 mm diameter and 300 mm length, and they con-

formed to [36]. The beam moulds were of 150 mm

square cross-section and 700 mm length, confirming

to [36]. The total number of cylinder specimens as

Table 2 Mix design with clay as deleterious material

Mix designa-

tion

Percent-

age of

clay with

respect to

sand

Quantities of materials per cubic meter of concrete

Re-

marks Ce-

ment

(kg)

Wa-

ter

(L)

Plasticiz-

ers.

(kg)

Sand

(kg)

Deleteri-

ous mate-

rial

(kg)

Coarse aggre-

gate (kg)

20mm 10mm

CM Nil 334 160 2 806 Nil. 624 416 Control

Mix

C1 0.25 334 160 2 804 2 624 416 -

C2 0.50 334 160 2 802 4 624 416 -

C3 1.00 334 160 2 798 8 624 416 IS Code

Limit

C4 1.50 334 160 2 794 12 624 416 -

C5 2.00 334 160 2 790 16 624 416 -


15

well as the total number of beam specimens was 18

(Table 3). All the joints between the various sections

of moulds were thinly coated with mould oil in order

to ensure that no water escapes during the casting

procedure as well as to prevent adhesion of the con-

crete to the moulds.

2.4.5 Compaction

The concrete was filled into the mould in layers

approximately 5 cm deep. In placing each scoopful

of concrete, the scoop is moved around the top edge

of the mould. A total of 35 strokes were given with

tamping rod for each layer thus laid. The strokes

were given in a manner such that the tamping rod

penetrated all the underlying layers. The sides of

mould were also tapped to close the voids, which

might have been left by tamping rod. Thus, dense

compact concrete samples were casted for later test-

ing.

2.4.6 Specimen identification

After the initial setting, all the specimens (cu-

bes, cylinders and beams) were labelled with perma-

nent marker with mix designation as show in Table

3 along with the date of casting.

2.4.7 Curing

The test specimens, in their respective moulds,

were stored in a place free from vibration and away

from direct sunlight for 24 ± 1/2 hours from the time

of mixing. After this period, the specimens were de-

moulded, marked and submerged in clean, fresh wa-

ter for the purpose of curing. The water in which the

specimens were submerged was periodically re-

newed, every seven days.

2.5 Standard procedure for tests on concrete

specimens

The samples after casting are shown in Fig. 3(a)

for cube, Fig. 3(b) for cylinder, and Fig. 3(c) for

beam. At the designated date of testing, the samples

were taken out of the curing pond and tested accord-

ing to the standard procedures, discussed in subse-

quent sub-sections.

2.5.1 Slump test for workability

Slump cone method of testing was used for

evaluation of workability according to the Indian na-

tional standard [31]. The internal surface of the

mould was cleaned and made free from superfluous

moisture and any set concrete before commencing

the test. The mould was placed on a smooth, horizon-

tal, rigid and non-absorbent surface, in the form of a

carefully levelled metal plate. The mould was firmly

held in place while it was filled. The mould was

filled in four layers, each approximately one-quarter

of the height of the mould. Each layer was com-

pacted with 25 strokes of the rounded end of the

tamping rod, distributed in a uniform manner over

the cross-section of the mould, which for the second

and subsequent layers penetrated into the underlying

layer/s. The bottom layer was tamped throughout its

depth. After completely filling the mould, the con-

crete was struck off level at the top with a trowel, so

that the mould is exact1y filled. Any mortar which

might have leaked out between the mould and the

base plate was cleaned away. Then, the mould was

removed from the concrete by raising it slowly and

carefully in a vertical direction. This allowed the

concrete to subside and the slump was measured im-

mediately thereafter by determining the difference

between the height of the mould and that of the high-

est point of the particular specimen. The slump

measured was recorded in terms of millimetres of

subsidence of the specimen during the test.

Table 0 Test matrix

Mix Designation

Number of cubes casted for compressive strength

test

Nos. of cylinders

casted for split

tensile test

Nos. of beams

casted for flex-

ural strength

test

7 days 14 days 28 days 63 days 91 days 28 days 28 days

CM 3 3 13 3 3 3 3

C1 3 3 13 3 3 3 3

C2 3 3 13 3 3 3 3

C3 3 3 13 3 3 3 3

C4 3 3 13 3 3 3 3

C5 3 3 13 3 3 3 3


16

2.5.2 Cube test of compressive strength

Specimens were tested immediately on removal

from the water and while they were still in the wet

condition [34]. The bearing surfaces of the testing

machine was wiped clean and any loose sand or other

material were removed from the surfaces of the spec-

imen which would be in contact with the compres-

sion plates. The specimen was placed in the machine

such that the load would be applied to opposite sides

of the cubes as cast, that is, not to the top or bottom.

The axis of the specimen was carefully aligned with

the centre of thrust of the seating plates. The load

was applied without shock and increased continu-

ously at a rate of approximately 140 kg/cm2/min un-

til the resistance of the specimen to the increasing

load breaks down and no greater load can be sus-

tained [34]. The maximum load applied to the speci-

men was then recorded. The measured compressive

strength of the specimen was calculated by dividing

the maximum load applied to the specimen during

the test by the cross-sectional area, calculated from

the mean dimensions of the section. In general, In-

dian standard stipulates that an average of three spec-

imen values should be taken as the representative of

the sample, provided the individual variation is not

more than ± 15 percent of the average [1]. For 28

days' strength, 13 cubes were tested and the mean of

the 13 was taken as the representative compressive

strength at 28 days. For all other ages, the average of

three samples was adopted as the representative

compressive strength.

2.5.3 Split tensile test

The split tensile strength test on cylindrical

specimens was performed according to the guide-

lines of Indian national standard [35]. The test spec-

imen was tested in wet condition soon after its re-

moval from the curing tank. The surface in contact

with the loading plate was made smooth and free

from protruded aggregates or any kind of undula-

tions. The centre line was marked across length of

the cylindrical specimen on the opposite sides and

the dimensions of the specimen were measured, to

the nearest 0.2 mm. The weight of the specimen was

measured and recorded. The specimen was placed in

the loading machine between the edges of two angles

ISA 65x65x6. The specimens were so placed that the

edge of the angles coincides with the centre line

made on the opposite faces of the cylindrical speci-

men. The jig was then placed in the machine so that

the specimen was located centrally. Care was taken

so that the loading through the packing strip was

truly axial. The load was applied without shock and

increased continuously at a nominal rate within the

range 1.2 N/mm2/min to 2.4 N/mm2/min [35]. The

maximum load at which the specimen failed was rec-

orded and the split tensile strength fst, of the speci-

men was calculated to the nearest 0.05 MPa using

Eq. (1) [35]:

𝑓𝑠𝑡 =2𝑃

𝜋𝑙𝑑 (1)

where,

fst = Split tensile strength of cylindrical specimen

(MPa)

P = maximum load applied to the specimen (N),

l = Length of the specimen (mm).

d = Cross sectional dimension of the specimen (mm).

2.5.4 Flexural test

The flexural strength test on beam specimens

was performed along the guidelines of Indian na-

tional standard [35]. As for the earlier samples, the

beam specimen was tested in wet condition, soon af-

ter its removal from the curing tank and the test was

conducted according to the Indian standard [34]. The

dimensions and weight of the specimens were rec-

orded before the testing. In this case, surface prepa-

ration was not required before placing the concrete

beam on the testing machine. The axis of the speci-

men was carefully aligned with the axis of the load-

ing device. The test setup was done on normal com-

pressive strength testing machine. A channel of

depth 150 mm was placed on the bottom loading

plate. Over it two angles were placed and welded

such that their vertex edge lied towards the top. The

distance between the two angles was maintained

such that their vertex was 600 mm apart. The beam

specimen was placed on these two angles without

disturbing the bottom channel and angle setup, with

50 mm overhang on either side of the vertices of the

angles. Another channel and angle setup, as de-

scribed above, was made for the top. In this case, the

distance between the vertices of the angles was kept

as 200 mm, and these would be the two-point loading

locations. When this was placed above the beam cen-

trally, it was thus possible to divide the beam into

three equal parts of 200 mm each. After fixing the

setup the load application was started, without shock

and increasing continuously at a rate such that the

extreme fibre stress increased at approximately 7

kg/cm2/min, that is, at a rate of loading of 400 kg/min

for the 15·0 cm specimen [34]. The load was in-

creased until the specimen failed, and the maximum

load was recorded. The appearance of the fractured

faces of concrete and any unusual features in the type

of failure was also noted. The flexural strength of the

specimen was expressed as the modulus of rupture

fcr, which, if ‘a’ equals the distance between the line

of fracture and the nearer support, measured on the

centreline of the tensile side of the specimen, in cm,

shall be calculated to the nearest 0.5 kg/sq.cm using

Eq. (2) when ‘a’ is greater than 20.0 cm for 15.0 cm


17

specimen, or Eq. (3) when ‘a’ is less than 20.0 cm

but greater than 17.0 cm for 15.0 cm specimen, ac-

cording to the failure condition:

𝑓𝑐𝑟 =𝑝𝑙

𝑏𝑑2 (2)

𝑓𝑐𝑟 =3𝑝𝑎

𝑏𝑑2 (3)

Where

fcr = flexural modulus of concrete (MPa)

b = measured width of the specimen (mm),

d = measured depth in mm of the specimen at the

point of failure (mm),

l = length in mm of the span on which the specimen

was supported (mm), and

p = maximum load applied to the specimen (N).

3. Results and discussion

In this section, the results of the experiments on

workability of green concrete and the strength of the

hardened concrete are presented and the salient ob-

servations are discussed.

3.1 Effect on fresh concrete: workability using

slump test

The results of slump value for concrete mix with

clay fines as deleterious material are presented in Ta-

ble 4. The effect observed on the slump values at the

explored concentrations (0.25% to 2%) of clay as

deleterious material in fine aggregate is reduction

ranging from 7% to 22%. The clay present in the

concrete mix absorbs the water and reduces the free

water thereby reducing the workability. This might

also affect the strength properties due to possible

lack of water for the purposes of complete hydration

of cement. The prevalent practice of evaluation of

strength at constant workability (by suitable altera-

tion of water-cement ratio or by addition of suitable

admixtures), as reported in literature, would be lim-

ited in examining this aspect. This study targets to

identify such possible effects on the strength proper-

ties of concrete by examining the cases with the de-

sign water-cement ratio and the admixture quantity.

The variations in the slump has been presented

in pictorial form in Fig. 4, where it can be clearly

noted that even at the IS code limit of 1% (C3) the

slump is lower than the target slump of 120 mm and

hence the IS code [3] limit of clay of 1% is unac-

ceptable for this mix. Thus, for the present case, the

limit of clay in concrete is recommended to be re-

vised to 0.5% according to the findings of this study.

In literature [30], it had earlier been indicated that the

stipulations in codes regarding the microfines such

as clay might prove inadequate depending upon their

(a) (b) (c)

Fig 3: Samples (a) cube; (b) cylinder; (c) beam;

Table 4 Results of slump test on green concrete with clay as deleterious material

Mix designation

Deleterious

material (clay)

(%)

Slump (mm) Percent variation Remarks

CM Nil 135 - Control Mix

C1 0.25 125 7 Acceptable slump

C2 0.50 120 11 Target slump

C3 1.00 110 19 IS code limit - Unac-

ceptable slump

C4 1.50 110 19 Unacceptable slump

C5 2.00 105 22 Unacceptable slump


18

mineralogical compositions and this was proved

true in this case. Further, that the presence of small

amount of clay in concrete might reduce workability

to a significant extent, as inferred from this study,

was earlier reported in literature [10].

3.2 Effect on hardened concrete: compressive

strength

The summary of the average compressive

strength of cubes for the different ages and the vari-

ous percentages of clay (deleterious material) are

presented in Table 5 (detailed results are provided in

the Appendix A). The general trend is observed that

for all ages, the higher percentages of deleterious

material (clay) result in lower strength. The standard

deviation obtained from the 13 samples (as elabo-

rated in Appendix A) for 28 days compressive

strength are included in Table 5. The small values of

standard deviation (between 1.04 MPa and 1.57

MPa) indicate that the quality control implemented

for the experiment was good. However, it is to be

noted that the standard deviation estimated from

small samples might be prone to errors, and should

be used with caution.

The compressive strength as a function of in-

creasing clay percentages in concrete are plotted in

Fig. 5a for the various ages. It was observed that at

the IS code [3] limit (1%), the compressive strength

obtained with the design mix attains the target com-

pressive strength of 31.6 MPa. In fact, the target

compressive strength could be achieved with inclu-

sion of up to around 1.6% clay. Hence from com-

pressive strength considerations, the IS code limit

may be concluded to be conservative. It may be

noted that at the IS code limit of 1% (C3), the com-

pressive strength reduced by around 14%, compared

to the control mix (CM), whereas the reduction in-

creased to around 32% for 2% clay (C5). The

strength reduction with increasing percentage of clay

appears to be linear in this case for all ages, except

the 7-day tests.

Fig 4: Variation of workability: slump

Table 5 Average compressive strength (MPa) of hardened concrete with clay as deleterious material

Mix designation

Age of concrete (days) Percent

variation

of 28

days

Strength

Standard

deviation of

28-days

Strength

7 days 14 days 28 days 63 days 91 days

Remarks:

28-day

strength

CM 31.08 38.27 41.78 47.45 48.71 - 1.57 Acceptable

C1 32.07 36.77 39.73 43.94 46.39 - 5 1.25 Acceptable

C2 31.88 35.96 37.28 42.34 44.15 - 11 1.04 Acceptable

C3 30.07 32.95 35.80 39.69 41.35 - 14 1.07 Acceptable

C4 27.14 30.79 33.30 36.34 38.25 - 20 1.26 Acceptable

C5 23.13 25.91 28.34 31.30 33.04 - 32 1.30 Unaccepta-

ble


19

The rate of gain of strength appeared to be af-

fected, particularly in the initial period of hydration.

This aspect is examined in the Fig. 5b where the

strength development is plotted against the age, for

the different mixes. It can be seen in Fig. 5b that the

slope of strength gain between age 7 and 14 days is

steeper for Mix CM than the mixes with clay as del-

eterious material, which indicates that the presence

of clay reduces the rate of strength gain in concrete.

From 7 days to 28 days, the strength gain for Mix

CM is 34.4%, for Mix C1 it is 24% and for remaining

mixes the gain is around 20%–with reference to the

strength at 7 days. This shows that rate of strength

gain or the rate hydration is affected due to presence

of clay fines. This may be attributed to the high water

absorbing property of clay minerals. From 7 days to

28 days, the rate of gain for all mixes appears to be

similar and a considerable strength gain takes place.

Increase in average compressive strength con-

tinued beyond 28 days, and was observed to be

around 15% from 28 to 91 days. Use of fly-ash-based

cement, as in this case, improves the concrete

strength at later ages due to the pozzolanic reaction

[5]. As noted earlier the final compressive strength

at any age for concretes with higher percentage of

clay fines is lower [17]. This would be possibly

(a)

(b)

Fig 5: Variation of compressive strength: (a) with clay percentage (b) with age


20

because the presence of clay fines in cement paste

matrix. Clay fines are soft, have high affinity to wa-

ter and exhibit high swelling properties [17,18]. This

swelled clay particles would dry up with progress of

hydration of cement and time, leaving several voids

within concrete–which could lead to reduced

strength. The presence of clay in the interfacial tran-

sition zone may further contribute towards the re-

duced mechanical properties of concrete [18].

3.3 Effect on hardened concrete: tensile strength

using split tensile test

The results of the split tensile test of concrete

mix with clay fines as deleterious materials are pre-

sented in Fig. 6 and the details are tabulated in Table

6. As mentioned earlier, the split tensile strength test

was conducted with concrete cylinders at 28 days

age of concrete, for all the mixes. It is generally ob-

served that the tensile strength evaluated from split

tensile test reduced with the increasing clay percent-

ages in concrete, and the variation was non-linear in

the range investigated.

From Table 6, it may be noted that at the IS code

limit of 1% (C3), the tensile strength reduced by

around 17%, compared to the control mix (CM),

whereas the reduction increased to around 30% for

2% clay (C5). The percent reduction in tensile

strength at 28 days, evaluated from split tensile test

(Table 6, last column), for various percentages of

clay as deleterious materials are similar to the varia-

tions observed in compressive strength at 28 days

(Table 5, last column). The Indian standard [1] stip-

ulates the tensile strength should be determined from

the tests conducted on the specimens, and provide a

guideline for the flexural tensile strength (= 0.7 × √

fck MPa), but there is no suggested guideline for split

tensile strength. Hence, using a conversion factor of

0.8 from cube to cylinder compressive strength

[5,34], the correlation between the average cube

compressive strength (fcube) and the split tensile

strength was taken from literature as: 0.2585 ×

(fcube)(2/3) [37] and 0.1711 × (fcube)

(0.7) [38].

The split tensile strength of concrete as a func-

tion of the average cube compressive strength

worked out from these relationships are compared to

the experimental split tensile strength in Table 7. In

Fig 6: Variation of split tensile strength

Table 6 Split tensile strength of concrete with clay as deleterious materials

Mix desig-

nation

Date of

casting

Date of test-

ing

Split tensile strength (MPa) Average split

tensile strength

(MPa)

Percent

variation 1 2 3

CM 04-04-2019 02-05-2019 3.777 3.579 3.650 3.669 -

C1 08-05-2019 05-06-2019 3.282 3.438 3.310 3.343 -9

C2 09-05-2019 06-06-2019 3.084 3.155 3.268 3.169 -14

C3 16-05-2019 13-06-2019 2.943 2.985 3.169 3.032 -17

C4 20-05-2019 17-06-2019 2.900 2.759 2.985 2.881 -21

C5 20-05-2019 17-06-2019 2.504 2.631 2.603 2.579 -30


21

Table 7 Comparison of experimental split tensile strength of concrete with literature

Mix designa-

tion

Average cube

compressive

strength (MPa)

Split tensile strength from literature (MPa) Average split tensile

strength (MPa) [38] [37]

CM 41.78 2.33 3.11 3.669

C1 39.73 2.25 3.01 3.343

C2 37.28 2.15 2.89 3.169

C3 35.8 2.09 2.81 3.032

C4 33.3 1.99 2.68 2.881

C5 28.34 1.78 2.40 2.579

this case, it would appear that the relationships from

literature would yield conservative tensile strength

value, even up to 2% of deleterious materials (clay)

in concrete mix. However, determination of concrete

properties for a mix from actual tests should be fa-

vored over the use of values stipulated in literature,

to get more accurate estimates of the tensile strength.

3.4 Effect on hardened concrete: tensile strength

using flexural test

The results of the flexural test of concrete beams

with clay fines as deleterious materials are presented

in Fig. 7 and the details are tabulated in Table 8. As

mentioned earlier, the flexural test was conducted

with concrete beams at 28 days age of concrete, for

all the mixes. It is generally observed that the tensile

strength evaluated from flexural tests reduced with

the increasing clay percentages in concrete in an al-

most linear manner.

From Table 8, it may be noted that at the IS code

limit of 1% (C3), the tensile strength reduced by

around 12%, compared to the control mix (CM),

whereas the reduction increased to around 22% for

2% clay (C5). The percent reduction in tensile

strength at 28 days, evaluated from flexural test (Ta-

ble 8, last column), for various percentages of clay

as deleterious materials are lower than the variations

observed in both compressive strength at 28 days

(Table 5, last column) and tensile strength at 28 days

from split tensile test (Table 6, last column), though

the general reducing trend is common to all. As men-

tioned earlier, the Indian standard [1] stipulates the

tensile strength should be determined from the tests

conducted on the specimens, and provides a guide-

line for the flexural tensile strength (= 0.7 × √ fck

MPa) from the characteristic strength of concrete

(fck). The suggested values in international stand-

ards range from 0.623 × √ fc MPa to 0.6268 × √ fc

MPa [39,40], where the average cylinder strength

(fc) is used. The comparison is presented with the IS

code [1] in this study.

The flexural tensile strength of concrete as a

function of the characteristic compressive strength

(25 MPa) works out to be 3.5 MPa [1]. In this case,

it would appear that the IS guideline [1] would yield

conservative flexural tensile strength value for con-

crete, up to 2% of deleterious material (clay) in con-

crete mix. However, determination of concrete prop-

erties for a mix from actual tests should be favored

over the use of values stipulated in literature, to get

more accurate estimates of the tensile strength.

4. Summary and conclusions

Concrete is extensively used in infrastructure

development in modern times. Various replacement

options for aggregates for production of sustainable

concrete are being attempted and practiced, with a

thrust on use of locally available materials for the re-

placement. The use of 100% crushed fine aggregates

in concrete mix has been recently approved by Indian

standards, and for wider acceptance, particularly in

government sector, evaluation of the various proper-

ties of green and hardened concrete using crushed

aggregates has become relevant. Particularly, as the

crushed aggregates are prone to contain various del-

eterious materials such as crushed fines (or crusher

dust) and clay (from clay matrix in the rock from

which the aggregates are produced), evaluation of

their effects on properties of concrete needs detail in-

vestigation. In fact the codes and standards [3] allow

a certain limiting amount of deleterious materials in

concrete mix but are silent over the quantification of

their effects on the properties of concrete. In litera-

ture, various studies are reported wherein the effects

of deleterious materials were investigated, but in al-

most all the cases, the workability was kept constant

by addition of extra water or admixtures – which

might not be true in actual practice. The present ex-

perimental study targeted quantification of the effect

of clay fines in concrete mix, at and around the lim-

iting amount of clay fines according to the Indian

standard [3,4]. The novelty of the study over the ma-

jority reported in literature is that here the original

concrete mix proportions were maintained while

adding various percentages of clay as deleterious


22

Table 8 Tensile strength of concrete from flexural test with clay as deleterious materials

Mix desig-

nation

Date of

casting

Date of test-

ing

Split tensile strength (MPa) Average split

tensile strength

(MPa)

Percent

variation 1 2 3

CM 04-04-2019 02-05-2019 4.764 5.031 4.907 4.901 -

C1 08-05-2019 05-06-2019 4.782 4.711 4.889 4.794 -2

C2 09-05-2019 06-06-2019 4.640 4.676 4.569 4.628 -6

C3 16-05-2019 13-06-2019 4.409 4.320 4.249 4.326 -12

C4 20-05-2019 17-06-2019 3.947 4.142 4.213 4.101 -16

C5 20-05-2019 17-06-2019 3.929 3.822 3.787 3.846 -22

material in concrete mixes. The experiments in-

cluded the slump test for workability; compressive

test, split tensile test and flexural test for strength

properties of concrete mix with deleterious materi-

als. The conclusions from the study are as follows:

1) The workability of concrete measured by

slump test reduced with increase in percentage of

clay from 7% reduction at 0.25% to 22% reduction

at 2%. This would be due to the higher water absorp-

tion by the clay particles, thereby reducing the free

water and resulting in harshness of mix. The ob-

served slump equaled the target slump at 0.5% of

clay and was less than target slump for all higher clay

percentages in concrete. Hence, the limiting clay per-

centage is suggested to be fixed at 0.5%, as against

the IS code [3] stipulation of 1% of fine aggregates.

2) The 7-day compressive strength increased

slightly for clay up to 0.5%, and reduced thereafter.

For all other ages, the compressive strength reduced

with increased clay percentages monotonically. The

reduction varied from 5% for 0.25% to 32% for 2%

clay in concrete mix. This could be attributed to the

swelling of clay with absorption of water [19] and

creation of voids due to drying of the same, thereby

reducing strength [17, 19]. Another possible reason

could be weakening of the interfacial transition

zones due to presence of clay [17]. For 2% clay fines

the average 28 day compressive strength was below

the target strength and thus would be unacceptable.

However, the IS code [3] stipulated value of 1% clay

fines would result in acceptable target strength.

3) The split tensile strength reduced with in-

creasing clay percentages monotonically, but in a

non-linear fashion. The reduction varied from 9% for

0.25% to 30% for 2% clay in concrete mix. There is

no guideline in IS code for the split tensile strength

and so the experimental results were compared to

values suggested in literature [37,38]. The split ten-

sile strength obtained for the concrete mixes with up

to 2% of clay as deleterious material was higher than

the values stipulated in literature [37,38].

4) The flexural tensile strength or modulus of

rupture reduced with increasing clay percentages

monotonically, apparently in a linear fashion. The re-

duction varied from 2% for 0.25% to 22% for 2%

clay in concrete mix. The flexural tensile strength

obtained for the concrete mixes with up to 2% of clay

as deleterious material was higher than the values

stipulated in IS code [1].

The major findings of this study regarding re-

duction of workability of concrete due to presence of

small amounts of clay are similar to [10], and that the

limits on microfines such as clay stipulated by the

codes might fall short of desired properties and

would have to be revised was indicated by [30]. The

findings that there was no strength improvement

(none except 7 days compressive strength up to

around 0.75% of clay), only reduction, was reported

in the same study earlier [30]. Increase in strength of

concrete due to addition of clay as reported exten-

sively in literature [9,10,13,15,29] was not observed

in this case, for any of the strength parameters exam-

ined.

Further studies are under progress for evalua-

tion and quantification of the effects of crusher fines

for the concrete mix for this site. It is recommended

that when crushed aggregates are being used for con-

crete, the effects of the possible deleterious materials

on the properties of green (workability) and hard-

ened (compressive and tensile strength; durability)

concrete should be carefully investigated, without

addition of extra water or admixtures for maintaining

constant workability. As the deleterious materials

would be site specific, separate investigation for

each source of the aggregates would be desirable.

Future studies might be towards investigating the ef-

fects of other deleterious materials or even various

combinations of the same on the workability,

strength and durability properties of concrete.

Acknowledgements

The first author sincerely acknowledges the en-

couragement and gracious support received from Dr.

M. Rajashaker (EIC, NFC, Kota) and Shri. P. A.


23

Pratap (Project Director, NFC, Kota) during this re-

search work. The excellent contributions from the

Tata Projects Ltd. Team for providing the laboratory

infrastructure and generous help in conducting the

experiments are gratefully appreciated.

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https://doi.org/10.1016/j.jobe.2018.04.022

26. Arulrajah, A.; Ali, M.M.Y.; Piratheepan, J.; and

Bo, M.W. (2012) "Geotechnical Properties of

Waste Excavation Rock in Pavement Subbase

Applications," Journal of Materials in Civil

Engineering, ASCE, 24(7), pp. 924~932. ).

https://doi.org/10.1061/(ASCE)MT.1943-

5533.0000419

27. VanWyk P.R.; and Croucamp, L. (2014)

"Evaluation of rock types for concrete aggregate

suitability for the construction of a gravimeter

vault and access road at the Matjiesfontein

Geodesy Observatory site near Matjiesfontein,

South Africa," Journal of South African

Institution of Civil Engineering, 56(2), 30~36.

28. Magnusson, S.; Lundberg, K.; Svedberg, B.; and

Knutsson, S. (2015) "Sustainable management

of excavated soil and rock in urban areas - A

literature review," Journal of Cleaner

Production, 93, 18~25.

29. Okamura, H.; and Ouchi, M. (2003) “Self

Compacting Concrete,” Journal of Advanced

Concrete Technology, 1(1), pp. 5~15.

30. Cramer, S. (2011) Defining the Impact of

Aggregate Fine Particles on Concrete Pavement

Performance, Wisconsin Highway Research

Program Report no. 0092-07-02, Department of

Transportation, Wisconsin, USA.

31. IS 1199: 1959 (Reaffirmed 2004), Method of

Sampling and Analysis of Concrete, Bureau of

Indian Standard, India.

32. Bhardwaj, B.D. (1970) Upper Vindhyan

Sedimentation in the Kota-Rawatbhata Area,

Rajasthan, PhD Thesis. Aligarh Muslim

University, Aligarh, India.

33. IS 1498: 1970 (Reaffirmed 2007), Classification

and Identification of soils for General

Engineering Purposes, Bureau of Indian

Standard, India.

34. IS 516: 1959 (Reaffirmed 2004), Methods of

Tests for Strength of Concrete, Bureau of Indian

Standard, India.

35. IS 5816: 1999 (Reaffirmed 2004), Splitting

Tensile Strength of Concrete – Method of Test,

Bureau of Indian Standard, India.

36. IS 10086: 1982 (Reaffirmed 2008),

Specification for Moulds for Use in Tests of

Cement and Concrete, Bureau of Indian

Standard, India.

37. Raphael, J.M. (1984) "Tensile Strength of

Concrete," ACI Materials Journal, 81(2), pp.

158~165.

38. Oluokun, F.A. (1991) "Prediction of Concrete

Tensile Strength from Compressive Strength:

Evaluation of Existing Relations for Normal

Weight Concrete," ACI Materials Journal, 88(3),

pp. 302~309.

39. ACI 435R-95 (1995: Reapproved 2000),

Control of Deflection in Concrete Structures,

Report by ACI Committee 435, American

Concrete Institute, USA.


25

40. ACI 318R-14 (2014) Building Code

Requirements for Structural Concrete, Report

by ACI Committee 318, American Concrete

Institute, USA.


Vol. 6, No. 1, pp. 26-36, June 2020

ISSN 2465-7964 / eISSN 2465-7972

https://doi.org/10.18702/acf.2020.6.6.26

26

Technical Paper

Optimization and evaluation of ultra high-performance con-

crete

P. N. Ojha, Piyush Mittal*, Abhishek Singh, Brijesh Singh, V. V. Arora

(Received February 17, 2020; Revised June 9, 2020; Accepted June 9, 2020; Published June 30, 2020)

Abstract: Ultra-High Performance Concrete (UHPC) has been defined as a cementitious based compo-

site material with compressive strength above 150 MPa and enhanced durability via its discontinuous

pore structure. The microstructure of UHPC is denser and more homogeneous in comparison to conven-

tional concrete. UHPC has several advantages over conventional concrete but the use of it is limited due

to the high cost and limited design codes. Methodology for production and development of UHPC needs

to be established. The paper covers both optimization and evaluation of Ultra-High-Performance Con-

crete along with highlighting the importance of packing density, mixing procedure and curing regimes

containing a high volume of mineral admixture and ultrafine materials. Cementitious content of all the

mixes in the study was kept in the range of 1000 kg/m3 and water to binder ratio was kept as 0.17. This

study focuses on the methodology to be adopted for optimizing the packing density of UHPC, the chal-

lenges associated with it and their influence on compressive strength.

Keywords UHPC; Packing Density; Compressive strength; Distribution Modulus; Curing Regime.

1. Introduction Over the last two decades, remarkable

advances have taken place in the research and

application of Ultra-High Performance Concrete

(UHPC) [1]. UHPC is the ‘future’ material with the

potential to be a viable solution for improving the

sustainability of buildings and other infrastructure

components. Ultra-high-performance concrete

(UHPC) with more than 150 MPa compressive

strength [2] generally incorporates relatively high

dosages of silica fume and superplasticizer, with a

relatively low concentration of aggregates of small

size. Some distinguishing features of UHPC include

an optimized gradation of the granular matter for

achieving high packing density, and water to

cementitious materials ratio of less than 0.25 [3]. The

hydrated paste in UHPC has a dense microstructure

which provides a distinct balance of strength, imper-

meability and durability [4, 5]. The superior mechan-

ical and durability characteristics of UHPC have led

to its use in the rehabilitation of concrete structures

[6, 7]. Recent developments in this field have em-

phasized the broadening of the raw materials selec-

tions and the use of common concrete production

methods to facilitate commercial applications of

UHPC [7, 8]. The basic principles for the develop-

ment of UHPC are as follows [9, 10]:

• Minimizing composite porosity by optimizing

the granular mixture through a wide distribution

of powder size classes and reducing the wa-

ter/binder ratio.

• Enhancement of the microstructure by the post

set heat treatment to speed up the pozzolanic re-

action of Silica Fume and other ultrafine ce-

mentitious materials to improve the mechanical

properties.

• The optimal usage of superplasticizer to reduce

water/binder ratio and improve workability.

• Improvement of homogeneity by eliminating the

coarse aggregate resulting in a decrease in me-

chanical effects of heterogeneity.

It has been well recognized by many researchers

that increasing the packing density of the granular

mix could lead to a better performance of the con-

crete. Increasing the packing density of the granular

mix would decrease the volume of paste needed to

P. N. Ojha is a Joint director at National Council for Cement


Corresponding author Piyush Mittal is a Project Engineer at

National Council for Cement and Building Materials (NCB).

Abhishek Singh is a Project Engineer National Council for Ce-

ment and Building Materials (NCB).

Brijesh Singh is a Manager at National Council for Cement


V. V. Arora is a Joint Director at National Council for Cement

and Building Materials (NCB)


27

fill up the voids and increase the amount of ‘‘ex-

cess’’ paste that could be utilized to improve the

workability of the concrete. Drawing analogy to the

case of a paste, increasing the packing density of the

cementitious materials would decrease the volume of

water needed to fill up the voids and increase the

amount of ‘‘excess’’ water that could be utilized to

improve the flowability of the paste. Hence, the key

to the production of UHPC, which demands both a

low water/binder (w/b) ratio to be used and high

workability to be attained, is the maximization of the

packing density of the granular skeleton of the con-

crete [11,12]. The effectiveness of supplementary

cementitious material in filling up the voids or in im-

proving the packing of the cementitious materials is

dependent on the fineness of the supplementary ce-

mentitious material. In general, a broader range of

particle size distribution would yield a higher pack-

ing density. This is because, with a broader range of

particle size distribution, the medium particles would

fill up the voids between the large particles, the fine

particles would fill up the voids between the medium

particles and the very fine particles would fill up the

voids between the fine particles and so on, leading to

the removal of more voids by the successive filling

effect. The addition of a supplementary cementitious

material finer than Ordinary Portland Cement (OPC)

would broaden the range of particle size distribution

and thus increase the packing density. A supplemen-

tary cementitious material with a higher fineness is

more effective because it would produce a broader

range of particle size distribution. [13, 14, 15]. Parti-

cle packing density can be defined as the solid vol-

ume of particles in unit volume. It has been reported

that when water/cementitious materials ratio was re-

duced to as low as 0.14 by weight, concrete having

the strength of 165–236 MPa was produced [16]. By

maximizing the packing of all granular materials in

the concrete mix using the same packing model and

also applying other advanced production techniques,

concrete having strengths in the order of 200–800

MPa was developed [17]. In 1996, packing theory

was applied for the design of self-consolidating con-

crete and based on the successful outcome it was

concluded that the performance optimization of con-

crete is mainly a matter of improving the packing

density of its granular skeleton [18]. Particle optimi-

zation methods can be divided into three groups (fig-

ure 1):

Fig 1: Mix design approaches for optimizing packing density [13]

• Optimization curves: Groups of particles, with a

specific particle size distribution, are combined

in such a way that the total particle size distribu-

tion of the mixture is closest to an optimum

curve.

• Particle packing models: These models are ana-

lytical models that calculate the overall packing

density of a mixture based on the geometry of the

combined particle groups.

• Discrete element models: With numerical mod-

els, a ‘virtual’ particle structure from given par-

ticle size distribution is generated.

In the present study, approach of the ideal curve

has been adopted for the optimization of concrete


28

mix to attain the maximum possible packing density.

The fundamental work of Fuller and Thomsen

showed that the packing of concrete aggregates is af-

fecting the properties of the produced concrete [19].

They concluded that a geometric continuous grading

of the aggregates in the composed concrete mixture

can help to improve the concrete properties. Based

on the investigation of Fuller and Thompson [19, 20]

& Andreasen and Andersen, a minimal porosity can

be theoretically achieved by an optimal particle size

distribution (PSD) of all the applied particle materi-

als in the mix as per empirical equation 1 below:

P(D) = ( D

Dmax)

q…………… (1)

Where P (D) is a fraction of the total solids be-

ing smaller than the particle size D (μm), Dmax is the

maximum particle size (μm) and q is the distribution

modulus. However, in the above empirical equation,

the minimum particle size is not incorporated, while

in reality there must be a finite lower size limit. In

continuation of this study, Funk and Dinger pro-

posed a modified model based on the Andreasen and

Andersen Equation. In this study, all the concrete

mixtures are designed based on modified Andreasen

and Andersen model, which is as follows [20,21]:

𝑃(𝐷) =𝑑𝑞−𝑑𝑚𝑖𝑛

𝑞

𝑑𝑚𝑎𝑥𝑞

−𝑑𝑚𝑖𝑛𝑞 ….……… (2)

Where Dmin is the minimum particle size (μm).

Using the above particle packing model, UHPC

mixes were optimised and compressive strength of

about 150 MPa at 28 days with low binder content

was achieved. The modified Andreasen and Ander-

sen packing model has already been successfully em-

ployed in optimization algorithms for the design of

normal density concrete and Lightweight concrete

[9, 10].

Different types of concrete mixes can be de-

signed using above equation 2 by applying different

values of the distribution modulus q, as the value of

q influences the ratio between coarse and fine parti-

cles. Higher values of distribution modulus (q>0.5)

lead to coarser mixtures whereas smaller values

(q<0.25) results in mixes that will be rich in fine par-

ticles [20]. A very high value of q above 0.50 leads

to higher aggregates to paste volumetric ratio and

therefore, the lesser paste will be available to lubri-

cate the fine aggregate particles, which in turn will

have decreased workability. Few trials were con-

ducted with a value of q below 0.37 and in those tri-

als as the value of q was lower, the concrete mix was

found to be stiff and required high-efficiency water

reducers. Hence, to obtain a workable mix with

available water reducing admixture and well-graded

particle size distribution, the value of distribution

modulus q has been taken as 0.37 in this study, keep-

ing in view the findings of the studies done by the

past researchers [20, 22]. Elrahman et al. also men-

tioned the work of Andreasen et al, wherein it was

suggested to use the exponent q in the range of 0.35

to 0.50 because fine particles are not able to pack

similar to bigger particles [22].

Table 1 Physical properties of materials

S. No. Properties Cement G.G.B.S Flyash UFGGBS Silica Fume

1. Fineness(m2/kg) 323 400 310 2026 16701

2. Specific Gravity 3.15 2.93 2.28 2.88 2.28

Table 2 Chemical properties of materials

S.No. Properties Cement G.G.B.S Flyash UFGGBS Silica

Fume

1 Loss of Ignition (LOI), % 2.3 0.33 0.4 0.17 2.73

2 Silica (SiO2), % 20.71 34.41 60.95 33.05 85.03

3 Iron oxide (Fe2O3), % 4.08 1.18 5.7 0.58 -

4 Aluminium oxide (Al2O3) , % 5.15 18.45 26.67 20.40 -

5 Calcium oxide (CaO), % 59.96 36.46 2.08 33.14 -

6 Magnesium oxide (MgO), % 4.57 7.00 0.69 7.62 -

7 Sulphate (SO3), % 1.84 0.097 0.29 0.19 -

8 Na2O, % 0.42 0.30 0.06 0.19 0.73

9 K2O, % 0.56 0.37 1.46 0.58 2.96

10 Chloride, % 0.012 0.022 0.009 0.016 -

11 Insoluble Residue, % 1.25 0.40 - 0.86 -

2. Materials

Properties of UHPC are highly dependent on the

type of material used in its production. In this study

cementitious material were selected in such a way

that particle size distribution has a wide range that

leads to the higher packing density of the concrete.

Cementitious Materials used in the study are OPC

53G, GGBS, ultrafine GGBS and Silica fume, Nano-

silica. Similarly, to increase the packing density of


29

aggregates, three different types of aggregates

namely Fine Quartz sand, Ground Quartz and Coarse

Quartz sand were used.

2.1 Cementitious materials

Cement: Ordinary Portland Cement (OPC) of

53 grade complying with IS 269: 2015 [23] was used

throughout the study. The physical properties of ce-

ment are tabulated in Table 1. Chemical properties

of cement have been listed in Table 2. Particles of

OPC 53 were in the range of 1.375 to 175 microns.

Particle size distribution has been shown in figure 5.

Ground granulated blast furnace slag

(GGBS): GGBS complying with IS 16714:2015

[24] was used in this study. Physical and chemical

properties have been listed in Table 1 and Table 2

respectively. Particles of GGBS were in the range of

1.15 to 250 microns. Particle size distribution has

been shown in figure 5.

Ultra-Fine Ground granulated blast furnace

slag (UFGGBS): Ultrafine ground powder of GGBS

used in this study has ultra-fineness which shows im-

provement in the properties of regular GGBS like

high surface area. UFGGBS is a specially processed

product based on high glass content with high reac-

tivity obtained through the process of controlled

granulation. For the production of UFGGBS, the

granulated slag with value-added material is ground

in a ball mill attached with a high-efficiency classi-

fier, which classifies the material and ensures that

only the required micro size particle enters the final

product. The entire process is operated by an auto-

matic process controller. Ultrafine GGBS commer-

cially available as Alccofine-1203 is a low calcium

silicate-based mineral additive which is generally

used as a replacement of silica fume in high-perfor-

mance concrete. Its latent hydraulic property and

pozzolanic reactivity results in the enhanced hydra-

tion process. UGGBS complying with IS 16715:

2015 [25] was used in this study. Physical and chem-

ical properties have been listed in Table 1 and Table

2 respectively. Particles of Ultra-Fine Ground gran-

ulated blast furnace slag are very fine and were found

to be in the range of 1 to 18 microns. Particle size

distribution has been shown in figure 5.

Silica Fume: Silica fume complying with IS

15388 [26] was used in this study. Physical and

chemical properties have been listed in Table 1 and

Table 2 respectively. Majority of the particles of Sil-

ica fume were found to be in the range of 0 to 10

microns. Particle size distribution has been shown in

fig.5.

Fly ash: Flyash complying with IS 3812 [27]

was used in this study. Physical and chemical prop-

erties have been listed in Table 1 and Table 2 respec-

tively. Particles of Flyash are slightly coarser than

the particle size of OPC and are in range of 1 to 300


figure 5.

Nano Silica: The mechanism of influence of

Nano silica on the cement hydration are available in

the literature. The studies indicated that the hydra-

tion heat of Ordinary Portland Cement blended with

Nano silica in the main period increases significantly

with an increased surface area of silica and the hy-

dration of tri-calcium silicate (C3S) gets accelerated

by the addition of Nano-scaled silica or CSH parti-

cles. A Nano silica slurry is selected as a pozzolanic

material to be used in this study. The solid content

and Brunauer–Emmett–Teller (BET) fineness are 20

(% w/w) and 24 m2/g. The specific density of Nano

silica was found to be 2.21 g/cm3.

2.2 Fine aggregates

Ground Quartz: Particle size of Ground

Quartz used in this project is on the coarser side of

the particle size of cement particles. It is used as a

micro filler to optimize the packing density of the

powder mix. Its particle size ranges from 0.5 to 140


figure 5. The microstructure of ground quartz by Op-

tical microscopy (Fig. 2) suggests that the minerals

present in order of abundance are quartz, orthoclase-

feldspar and iron oxide. Subhedral to anhedral quartz

grains with sharp grain margins are well-graded and

homogenously distributed. The majority of quartz

grains are in the size range of 20µm to 30µm. The

strained quartz percentage is about 9% and their un-

dulatory extinction angle (UEA) varies from 100 to

120. Subhedral orthoclase grains with smooth grain

margins are fresh in nature. Subhedral iron oxide

grains with sharp grain margins are randomly distrib-

uted in the sample.

Fig 2: Distribution of minerals grains in Ground

Quartz. (5x, X-Nicols)

Fine Quartz Sand: UHPC mixes were pro-

duced using quartz sand having a particle size rang-


30

ing from 150 to 996 microns. Particle size distribu-

tion has been shown in figure 5. The microstructure

of ground quartz sand by optical microscopy (Fig. 3)

suggests that it has a subhedral to anhedral quartz

grains with sharp grain margins which are well-

graded and homogenously distributed. The majority

of quartz grains are in the size range of 300µm to

600µm.

Fig 3: Distribution of minerals grains in fine Quartz

sand (11.25x)

Coarse Quartz sand: Coarse quartz sand used

in this study have a particle size ranging from 1mm

to 3mm. The microstructure of coarse quartz sand by

optical microscopy (Fig. 4) suggests that it has a sub-

hedral to anhedral quartz grains with sharp grain

margins which are well-graded and homogenously

distributed.

Fig 4: Distribution of mineral grains in Coarse

Quartz sand sample (11.25x)

Fig 5: Particle size distribution of materials

3. Experimental

3.1 Packing density

In the present research, the ideal curve method-

ology has been adopted. In this method, materials are

combined in such fractions that there combined grad-

ing lies close to a certain optimum curve given by

Modified Andreasen and Andersen equation as men-

tioned above in equation 2. The proportions of each

material in the mix are adjusted until an optimum fit

between the composed mix and the target curve is

reached, using an optimization algorithm based on

the Least Squares Method (LSM), as presented in

equation 3. When the deviation between the target

curve and the composed mix expressed by the sum

of the squares of the residuals (RSS) at defined par-

ticle sizes, is minimized, the composition of the con-

crete is treated as the most optimum one.

RSS = ∑ (𝑃𝑚𝑖𝑥(𝐷𝑖𝑖+1) − 𝑃𝑡𝑎𝑟(𝐷𝑖

𝑖+1))2𝑛

𝑖=1… (3)

where, Pmix is the composed mix and Ptar is the target

grading.

Around 40 mixes with cementitious materials of

OPC-53, GGBS, UFGGBS & Silica fume, Flyash

and aggregates including Fine Quartz Sand, Quartz

powder and Coarse Quartz sand were theoretically

optimized for optimum particle packing with the

help of above mentioned Modified Andreasen and

Andersen equation. The value of q adopted in this

research is 0.37 based on the literature. The mixes


31

with the least value of RSS were selected for experi-

mental study. Mixing of UHPC requires equip-

ment that provides more energy and shear than reg-

ular concrete mixers due to the low water content and

high powder content. In general, the expected perfor-

mance (including fresh and hard-solid properties) of

the selected mixture cannot be achieved when low-

energy mixers are used to mix UHPC. Moreover, the

use of a low-energy mixer will also increase the turn-

over time of the mixture, causing the temperature of

the mix to rise, which is detrimental to the UHPC

mixing process [28, 29]. Therefore, for homogene-

ous mixing of UHPC authors designed and devel-

oped planetary mixer of 60 litre capacity with high

mixing efficiency that can be operated at variable

speed with maximum speed up to 325 RPM. Selected

mixes were cast using the developed planetary

mixer. Compressive strength was measured using a

cube with a specimen size of 70.6 mm. Wet packing

density was also determined using the below-men-

tioned equation given by Kwan [30, 31, 32]. If a

UHPC mix consists of several different materials de-

noted by α, β, γ and so forth, ρ and R are solid density

and volumetric ratio of the respective material. uw is

the ratio of water to solid content. Then, the solid

volume of the cementitious materials Vc and wet

packing density ɸ may be worked out from equations

4 and 5 mentioned below:

𝑉𝑐 = 𝑀

ρ 𝑤u 𝑤+ρ αR α+ρ βR β+ρ γR γ …......... (4)

ɸ = 𝑉𝑐

𝑉................................ (5)

As per numerous researchers, the packing den-

sity of concrete is in the range as given below in table

3.

Table 3 Packing density of different types of concrete [27]

Type of Concrete Packing Density range

Normal strength Concrete 0.65 - 0.72

High Strength Concrete 0.72 – 0.8

Ultra high-performance concrete More than 0.8

3.2 Curing regimes

Heat treatment of concrete plays a significant

impact on the rate of strength development and also

has a strong influence on mechanical properties as

well as micro-texture of UHPC. Apart from the ef-

fect of accelerating the curing process and increasing

early strength, heat treatment could be useful for ce-

mentitious systems containing supplementary ce-

mentitious materials such as silica fume, granulated

blast furnace slag, fly ash by influencing the reaction

rate of these mineral additions compared to the same

system cured at ambient conditions. Compared to the

UHPC cured under ambient conditions, heat-treated

specimens of UHPC show generally a denser micro-

structure that can lead to an increase in compressive

strength and thus can improve overall mechanical

properties of UHPC [33, 34]. In the present study,

three different curing regimes are used as given be-

low:

a) Autoclave curing at 2.1 MPa and 215ºC for 8

hours followed by Standard curing till the age

of testing (up to 07 days)

b) Steam curing at 90ºC and 100% RH for 24

hours followed by Standard curing till the age

of testing (up to 28 days)

c) Standard water curing until the age of testing

(28 days).

3.3 Mix design details

In the present study, the proportion of individ-

ual cementitious material was decided using the

Modified Andreasen and Andersen equation. Out of

40 mixes that were theoretically optimized, three fi-

nal mixes were cast in the laboratory for a study on

compressive strength. The details of the composition

of the cementitious material of these three optimized

mixes are discussed later. The total cementitious

content in concrete mixes is kept around 1000 kg/m3.

To attain better particle packing density, a combina-

tion of three fine aggregates were used i.e. Ground

quartz, Fine quartz sand and Coarse quartz sand in

the proportion of 30:40:30 respectively. The nano-

silica was used as a 3% replacement to OPC content.

Dosage of PC based superplasticizer was kept as 2%

by weight of cementitious material.

3.4 Mix methodology

UHPC has been produced using a wide variety

of mixers, ranging from laboratory-sized pan mixers

to revolving drum truck mixers. In general, mixing

UHPC is a somewhat different process than mixing

conventional concrete. UHPC typically includes a

limited amount of water and no coarse aggregate. As

such, the UHPC requires the input of extra mixing

energy both to disperse the water and to overcome

the low internal mixing action from the lack of

coarse aggregate. Different researchers have adopted

various mixing protocols to achieve homogenization

of the UHPC mixture in a shorter span. Although


32

specific details of the overall mixing process dif-

fered, all researchers were unified in that UHPC

components had to be dry mixed before adding water

and superplasticizer. Dry mixing intends to ensure

higher bulk density and lower moisture require-

ments. A typical mixing process involves first charg-

ing the mixer with the dry components and ensuring

that the mix is blended homogenously. Thereafter,

water and chemical admixtures are added and dis-

persed. Mixing continues, sometimes for an ex-

tended period depending on the mixer energy input,

until the UHPC changes from a dry powder into a

fluid mixture. The mixing methodology adopted in

this study is given in Table 4.

Table 4 Mixing methodology to prepare concrete mixes in this study

Step No. Mixing methodology Duration

Step 1 Dry mixing of all cementitious material and aggregates at low speed of 125

rpm 3 minutes

Step 2 Adding 100% water & 75% superplasticizer and mix at medium speed of 250

rpm 5 minutes

Step 3 Add remaining superplasticizer and mix the constituents at high speed of 325

rpm 5 minutes

Step 4 Mixing continues until concrete achieve the required flow -

Table 5 (a) Details of proportion (% of total cementitious content) of several combinations of cementitious

materials

Mixes OPC 53 (%) Silica Fume (%) GGBS (%) UFGGBS (%) Flyash (%) RSS

M1 40 10 20 30 - 424.8

M2 40 10 10 40 - 465.3

M3 35 - 35 30 - 454.3

M4 20 15 35 30 - 434.0

M5 30 10 40 20 - 439.4

M6 70 10 - 10 10 442.2

M7 60 20 - 10 10 424.0

PM 1 60 10 20 10 - 418.0

PM 2 80 20 - - - 417.0

PM 3 60 10 - 10 20 380.0

Table 5 (b) Details of proportion (% of total fine aggregates) of several combinations of fine aggregates

Mixes Fine Quartz sand (%) Coarse Quartz sand (%) Ground Quartz (%) RSS

C1 70 10 20 437.3

C2 50 30 20 275.5

C3 30 50 20 291.7

C4 10 70 20 486.1

C5 50 10 40 340.9

C6 30 30 40 295.5

C7 10 50 40 428.1

C 8 60 10 30 367.5

C9 40 30 30 263.8

C10 20 50 30 338.2

4. Results & discussion

4.1 Optimization of mix

As mentioned earlier, in the present study the

ideal curve methodology was used for the optimiza-

tion of particle packing of concrete mix. For optimi-

zation of concrete mix, proportions of cementitious

materials and fine aggregates were optimized sepa-

rately. Several mixes were theoretically analyzed

with help of Modified Andreasen and Andersen

equation and mix with least RSS values were used

for laboratory study. Out of several sets of combina-

tions having different proportions, 10 sets of combi-

nations for cementitious materials and 10 sets of

combinations for fine aggregates with least RSS val-

ues have been tabulated below in table 5(a) and 5(b).

Dmin and Dmax (for optimizing the proportion of ce-

mentitious materials) used in modified Andersen and


33

Andreasen equation for the ideal curve are 0.112 mi-

crons and 1408 micron. Whereas, Dmin and Dmax (for

optimizing the proportion of inert fine aggregates)

used in modified Andersen and Andreasen equation

for the ideal curve are 0.5 microns and 2360 micron

respectively. Out of all the cementitious combina-

tions, PM1, PM2 and PM3 were selected to prepare

concrete. For fine aggregates, combination C9 was

selected to be used for all the three mixes (having

optimized cementitious combinations of PM1, PM2

and PM3) as it has the least RSS value in comparison

to other combinations. The combined particles size

distribution of selected cementitious combination

along with ideal particle size distribution calculated

using Modified Andreasen and Andersen equation is

given in Fig 6. The details of the composition of se-

lected concrete mixes are given in Table 6.

Table 6: Details of mix design

Fig 6: Grading of ideal curve and designed mixes

4.2 Compressive strength

As discussed earlier in section 4.1, cementitious

combinations, PM1, PM2 and PM3 with combina-

tion C9 of fine aggregates were selected to prepare

concrete. Concrete specimen of all the three mixes

were subjected to three curing regimes (mentioned in

3.2) and were evaluated for compressive strength.

For each mix, nine cubes were cast and an average

compressive strength of three cubes for each curing

regime has been plotted for these three selected

mixes (Fig 7). Apart from the interfacial transition

zone (ITZ) between aggregate and cement paste, the

role of the particle packing density is a major factor

governing the compressive strength of the UHPC.

For all three curing regimes studied, PM3 has a

highest compressive strength in comparison to PM1

and PM3. Wet packing density was determined as

per equation 4 and 5. It was observed that for all the

three mixes, value of wet packing density (ɸ) is more

than 0.8. Such a high value of wet packing density

suggests that mixes optimized using ideal curve

methodology have a dense microstructure that helps

in achieving ultra-high compressive strength. Com-

pressive strength in the case of the Autoclave curing

regime is significantly higher in comparison to the

other two curing regimes for all the three optimized

mixes. The maximum compressive strength

achieved is 186.5 MPa for mix PM3 in the case of an

autoclave curing regime. The increment in compres-

sive strength in case of steam curing varied from

12% to 47.3% and in the case of autoclave curing it

varied from 19.11% to 81.9%. Under the conditions

of high temperature and pressure, the chemistry of

hydration is substantially altered. CSH forms but is

converted to a crystalline product α -calcium silicate

hydrate (α-C2S) which causes an increase in porosity

and reduction in strength. However, in the presence

of silica α -C2S converts to tobermorite (C5S6H5) on

continued heating thus high strength can be obtained.

On the other hand, prolonged autoclaving may cause

the formation of other crystalline calcium silicate hy-

drates with a strength reduction. It is believed that

Mix Cement

kg/m3

Silica fume (SF) + GGBS (BS) +

UFGGBS (UFBS) + Flyash (FA)

+ Nano Silica (NS) kg/m3 W/B

Fine Quartz Sand (FQ)

+ Ground Quartz Sand

(GQ) + Coarse Quartz

Sand (CQ) kg/m3

Super

Plasticizer

kg/m3 SF BS UFBS FA NS

FQ GQ CQ

PM1 582 100 200 100 0 18 0.17 498.4 292.6 371.0 20

PM2 776 200 0 0 0 24 0.17 501.0 294.1 372.9 20

PM3 582.0 100 0 100 200 18 0.17 489.9 287.6 364.7 20


34

the complete conversion to tobermorite is not desir-

able and that there is an optimum ratio of amorphous

to crystalline material for maximum strength [35,

36]. The strength level after autoclaving generally

cannot be reached even with 24-hour steam curing

for all three mixes. This can be explained by the dif-

ferent hydration mechanisms due to these curing

methods. While steam curing increases the reactivity

of ingredients, autoclaving leads to the development

of different phases and incorporation of ultrafine ma-

terial is essential to achieve ultra-high compressive

strength.

Fig 7 Compressive strength for three different curing regimes

5. Conclusions

(1) In order to obtain a workable and satisfactory

mix with available water reducing admixture and

well-graded particle size distribution, the value

of distribution modulus q was adopted as 0.37.

Ideal curve methodology has been used for the

optimization of particle packing of concrete mix.

For the preparation of concrete mix, proportions

of cementitious materials and fine aggregates

were optimized separately. Several mixes were

theoretically analyzed and optimized with help

of Modified Andreasen and Andersen equation

and mixes with least RSS values were used for

laboratory study. Mixes optimized and prepared

using lower values of q may result in higher

compressive strength. However, such mixes

were found to be very stiff and required very

high-efficiency water reducers

(2) Wet packing density was determined as per the

method given by Kwan. It was observed that the

value of wet packing density (ɸ) for all the three

mixes is more than 0.8. Such a high value of wet

packing density suggests that mixes optimized

using ideal curve methodology have a dense mi-

crostructure that helps in achieving ultra-high

compressive strength.

(3) Cementitious content (fine particles) in a UHPC

mix is quite high in comparison to a normal

strength concrete mix. Mixing methodology and

type of mixer used for the preparation of UHPC

have a large influence on the mixing efficiency

and uniformity of UHPC which eventually af-

fects its properties in a fresh and hardened state.

The mixing methodology adopted in this study

and preparation of concrete mix using planetary

mixer developed for this study ensured homoge-

neous mixing without any lump formation.

(4) The curing regime plays a vital role in the devel-

opment of hardened properties of the UHPC

mix, which contains a high amount of mineral

admixtures. Post set heat treatment enhances the

microstructure by speeding up the pozzolanic re-

action of Silica fume and other ultrafine cementi-

tious materials to enhance the mechanical prop-

erties. From results, it can be observed that the

maximum compressive strength achieved was

186.5 MPa in case of autoclave curing for mix

PM3. The percentage improvement in compres-

sive strength in case of steam curing varied from

12% to 47.3% and in the case of autoclave curing

it varied from 19.11% to 81.9% in comparison to

the compressive strength observed in case of

standard curing.


35

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Cement and Concrete Research


Vol. 6, No. 1, pp. 37-51, June 2020

ISSN 2465-7964 / eISSN 2465-7972

https://doi.org/10.18702/acf.2020.6.6.37

37

Technical Paper

Bond and durability investigation of basalt fiber and PEN fi-

ber reinforced composites for concrete applications

Donguk Choi, Youngho Kim*, Batzaya Baasankuu, Gombosuren Chinzorigt

(Received September 23, 2019; Revised February 24, 2020; Accepted April 26, 2020; Published June 30, 2020)

Abstract: Bond and durability characteristics of basalt fiber reinforced polymer (BFRP) and polyeth-

ylene naphthalate (PEN) fiber/PEN FRP were investigated. Magnitude and distribution of the bond stress

between BFRP/PEN FRP and concrete were investigated by double lap shear test. Four different types

of durability test were performed: (1) Beam bond test following accelerated conditioning protocols by

ACI 440.9R using plain concrete beams strengthened with BFRP or PEN FRP; (2) tensile test of PEN

fiber/PEN FRP after immersion in 1N NaOH, 3% NaCl solutions, and water up to 6 months; (3) tensile

test of PEN fiber/PEN FRP after immersion in 5% and 10% diluted solutions of HCl; and (4) exposure

to natural outdoor environment. Bond test results indicated high bond stress developing over relatively

short distance for BFRP that has high elastic modulus (EBF = 68.4 GPa) while relatively low bond stress

developing over longer length for PEN FRP that has low elastic modulus (EPEN = 17.4 GPa). In the beam

bond test, very good behavior was shown by PEN FRP after 4 month’s exposure to wet and alkaline

conditions while moderate behavior was shown by BFRP. Overall, the performance of PEN fiber/FRP

was satisfactory in all durability tests conducted in this study.

Keywords: fiber reinforced polymer; double lap shear test; durability; beam bond test; basalt fiber;

PEN fiber.

1. Introduction

Various fibers, including carbon fiber (CF), ar-

amid fiber (AF), glass fiber (GF), and basalt fiber

(BF), are used for the purpose of external strength-

ening of RC structures and members. The mechani-

cal characteristics common to CF, AF, GF, and BF

include linearly-elastic stress-strain relationship,

high tensile strength, and high elastic modulus while

they have relatively small rupture strain (< 3%). On

the other hand, a new class of fiber including poly-

ethylene terephthalate (PET) fiber and polyethylene

naphthalate (PEN) fiber has non-linear stress-strain

relationship, good strength in tension, low elastic

modulus, and large strain capacity in tension (4%-

15%): i.e. PET/PEN fibers are often said to have

LRS (large rupture strain) capacity.

In this study, bond and durability characteristics

of two different fibers were investigated: BF and

PEN fiber. BF is an inorganic fiber produced from

natural basalt rocks by melting and extrusion process

[1, 2]. BF, which has been used in Civil Engineering

discipline only recently, is more economical than CF

or AF. Basalt fiber has excellent thermal resistance

such that it can be used as insulating material replac-

ing asbestos which poses health hazards by damag-

ing respiratory systems [1]. PEN fiber is a synthetic

fiber and is a product of petrochemical industry. PEN

fiber has good strength (over 800 MPa) and low elas-

tic modulus (about 1/4th of BF), exhibits non-linear

stress-strain behavior, and has large rupture strain in

tension (7%-9%).

Thin BFRP sheets are often used for external

strengthening in flexure and/or shear of RC beams

[3, 4]. BFRP sheet or the basalt fiber rope can also

be utilized for seismic strengthening of RC columns

in the form of external wrapping [5-7]. Bond charac-

teristics of BF bonded to concrete using adhesive

have been investigated [8-10]. Shen et al. studied

bond behavior of 21 test specimens by double lap

shear test subjected to different strain rate. Effective

bond length of BFRP ranged between 56-72 mm

while the effective bond length decreased with the

Donguk Choi is a Professor of School of Architecture and De-

sign Convergence of Hankyong National University, Korea.

Corresponding author Youngho Kim is a Professor of Depart-

ment of Chemical Engineering of Hankyong National Univer-

sity, Korea.

Batzaya Baasankuu is an Assistant Lecturer of Department of

Civil Engineering and Architecture of Mongolian University of

Science and Technology, Mongolia.

Gombosuren Chinzorigt is an Engineer at Gurvan Tahilgat

Orgil LLC, Mongolia.


38

increasing strain rate which varied between 70

mm/sec to 0.07 mm/sec [8]. Xie et al. studied the

bond behavior BFRP-concrete interface subjected to

wet-dry cycling in a marine environment using 26

small beams specimens [10]. The concrete beams

were externally strengthened using BFRP sheet on

the tension side. The test specimens were immersed

in salt water with concentration of 3.5% and sub-

jected to wet-dry cycles up to 360 days (one cycle

per day). All beam specimens were tested by 4-point

bending. For the control specimens, the failure mode

was debonding of the adhesive layer before exposure

while the failure mode changed to BFRP fracture in

tension after exposure to marine environment. The

fatigue strength of the BFRP/concrete interface after

the wet-dry cycling was approximately 60% of the

ultimate load capacity (of the control specimen). Ex-

isting durability test results of BFRP indicate that the

BF may also degrade under wet and alkaline condi-

tions [11]. The durability performance of BFRP can

be improved by coating the BF with zirconium diox-

ide or titanium dioxide [12-14].

Recent research explored possible application

of the LRS PET FRP and PEN FRP composites

mainly for the external strengthening of the RC col-

umns by confinement utilizing its large rupture strain

capacity [15-20]. Baasankhuu et al. compared the be-

haviour of concrete cylinders confined by BFRP and

PEN FRP [18]. The strength of confined concrete

wrapped by PEN FRP was achieved at much higher

axial/lateral strain of the concrete than the BFRP

wrapping: PEN FRP wrapped concrete deformed

more laterally to develop axial strength equivalent to

BFRP wrapped concrete. Park et al. reported results

of a rare study on the flexural strengthening of RC

beams using PET FRP [21]. Despite very low elastic

modulus (about 1/20th that of steel) of PET FRP, the

external strengthening by 6-mm-thick multi-layer

PET FRP sheet was effective to significantly im-

prove the flexural strength and ductility of the RC

beams. PET FRP sheet did not debond from the con-

crete substrate at ultimate.

At present, there are few studies that focused on

the bond behavior between LRS PEN FRP and con-

crete. In addition, few studies can be found in the ex-

isting literature either on the durability characteris-

tics of PEN FRP clearly indicating a research need.

One objective of this study was to investigate and

compare the bond behavior of BFRP-to-concrete and

PEN FRP-to-concrete interfaces in terms of magni-

tude and distribution of the bond stress. Bond behav-

ior was investigated by double lap shear test in this

study. The other objective was to investigate the du-

rability characteristics of BFRP, PEN fiber/PEN

FRP. Four different durability tests were performed:

(1) beam bond test by accelerated conditioning pro-

tocols (ACP) by ACI 440.9R for both BFRP and

PEN FRP; (2) durability test of PET fiber/FRP under

wet, saline, alkaline conditions up to 6 months; (3)

durability test of PET fiber/FRP under acidic condi-

tions; and (4) exposure to natural outdoor conditions.

The durability test program especially concentrated

on the PEN fiber/FRP.

Table 1 Mechanical properties of fiber roving and adhesive

Type Stress

(MPa)

Strain

(%)

E1

(GPa)

E2

(GPa)

Area

(mm2)

Thick-

ness

(mm)

Density

(g/mm3)

Basalt fiber roving 1226 1.95 68.4 n/a 0.45 0.113 0.0027

PEN fiber roving 822 8.03 17.4 8.30 2.00 0.840 0.0014

PEN fiber sheet 842 9.01 17.5 8.33 -- -- --

PEN FRP 912 9.13 21.4 8.59 -- -- --

Adhesive 40.9 2.58 1.59 -- -- -- --

NOTE: 1. Results show mean value of 12 tests for fibers and 5 tests for adhesive; 2. PEN fiber shows bilinear

response and so the slope of the first and the second line is shown as E1 and E2, respectively (see Fig. 1(b); 3.

Adhesive mechanical properties were tested 7 days after hardening.

2. Material properties and test method

2.1 Material properties of fibers and adhesive

Tensile properties of BF roving, PEN fiber rov-

ing, PEN fiber sheet, and PEN FRP (i.e. PEN fiber

sheet embedded by two-part epoxy) were tested fol-

lowing ISO 10406-2 with results summarized in Ta-

ble 1 [22]. Twelve tensile tests were completed for

each fiber/FRP while the mean values are shown in

Table 1. Figure 1(a) shows the stress-strain relation-

ship of BF and PEN fiber determined in this study.

In Figure 1(a), the stress-strain relationship is linear

for BF while it is non-linear for PEN fiber. The non-

linear stress-strain relationship of PEN fiber can be

modelled using a bilinear relationship with each line

having a slope of E1 (slope of the first line connecting

the origin and the stress corresponding to 1% strain)

and E2 (slope of the second line), respectively, as

shown in Table 1 and Fig. 1(b). For the double lap


39

shear test, beam bond test, and tensile test of the

BFRP and PEN FRP, a two-part epoxy (adhesive)

was used. The tensile properties of the adhesive were

determined following ASTM D638 [23]. As shown

in Table 1, tensile strength, ultimate strain in tension,

and elastic modulus of the adhesive are 40.9 MPa,

2.58%, and 1.59 GPa, respectively. A universal test-

ing machine (UTM) of 50-kN capacity was used for

tensile tests of fibers/FRP and adhesive.

(a) Stress-strain relationship

(b) Bi-linear model of PEN fiber

Fig. 1 Stress-strain relationship of BF and PEN

fibers

2.2 Double lap shear test

As the LRS PEN fiber has low elastic modulus,

a significant thickness of the PEN FRP is needed

when strengthening RC members such as RC beams,

columns, etc. [18]. Bond stress that develops at the

interface between the thick PEN FRP layer and the

concrete substrate can be high because a thicker FRP

may induce higher bond stress, which in turn may

lead to a premature debonding failure at the FRP-

concrete interface [24, 25]. Bond investigation by

double lap shear test was planned both for the PEN

FRP- and the BFRP-strengthened concrete speci-

mens to investigate magnitude and distribution of the

bond stress that develops between the FRP and the

concrete.

Modified double lap shear test setup was used

following recommendations of CSA S 806 Annex P

in general [26]. Figure 2 shows a double lap shear

test specimen used in this study. Size of the test spec-

imen was 140 (b) x 150 (h) x 550 mm (L). Concrete

with 30-MPa target strength was designed, cast in the

laboratory, and wet cured for 28 days. The 28-day

compressive strength was 36.0 MPa. A thin steel

plate placed at center of specimen at time of casting

practically disconnected a specimen into two con-

crete blocks of equal length. A 16-mm diameter

grooved steel rod was installed in the middle of each

concrete block in the axial direction with about 300-

mm length protruding outside the concrete block (for

use by testing grip of UTM). 28 days after casting,

two opposing side faces of the concrete blocks were

lightly roughened using a hand grinder. Two layers

of BF rovings or single layer of PEN fiber sheet was

applied (bFRP = 100 mm, LFRP = 500 mm), respec-

tively, on the roughened faces using the adhesive

with amount of 200% by vol. for BFRP and 150% by

vol. for PEN FRP (The same amount of adhesive was

used for BFRP and PEN FRP, respectively, through-

out this study). Seven days after application of the

adhesive, multiple 5-mm-long electronic strain

gauges were installed on the surface of BFRP and

PEN FRP starting from center of the specimen at 21-

mm spacing on center as shown in Fig. 2.

Table 2 Test variables of double lap shear test

Type No. of

layers

No. of

test

fck

(MPa)

Fiber thk.

(mm)

Adhesive thk.

(mm)

BFRP 2 2 36.0 0.226 0.452

PEN FRP 1 2 36.0 0.840 1.260

The double lap shear test specimen was sub-

jected to tensile force during test using 1,200-kN-ca-

pacity UTM operated in displacement control (ramp

rate = 1 mm/min). The applied load was measured

using load cell of the UTM. Signals from the strain

gauges and the load cell were electronically recorded

by a data logger. Two replicate specimens were

tested for each FRP type. Table 2 summarizes test

variables of the double lap shear test.


40

Fig. 2 Double lap shear test specimen [21]

2.3 Beam bond test

100 (b) x 100 (h) x 400 (L) mm plain concrete

beam specimens were used for the beam bond test

following guide to accelerated conditioning proto-

cols (ACP) of durability assessment by ACI 440.9R

[27]. Concrete with 60-MPa compressive strength

was cast in the laboratory and wet cured for 28 days.

A 50 x 100 mm acrylic plate (thickness = 2 mm) was

inserted in the beam mid-span on the tension side up

to beam half-height to create a notch as shown in Fig.

4(a). After 28 days of wet cure, two layers of BF

rovings or single PEN fiber sheet was externally

bonded on the bottom surface of the beam using ad-

hesive. Concrete surface was lightly roughened be-

fore the FRP application. Both BFRP and PEN FRP

were 80 mm wide and 200 mm long (See Fig. 4(a)).

A week after the FRP application, the specimens

were subjected to three different environmental con-

ditions: Room condition (T = 23°C +/- 3°C) and im-

mersion in water and 1N NaOH solutions, respec-

tively (Temperature of water and NaOH solution =

60°C +/- 3°C). After immersion for 3,000 hours (4-

months) in water and in the alkaline solution, the

beams were retrieved and stored in the room condi-

tion for two days. Flexure test by three-point loading

following recommendations of ACI 440.9R was car-

ried out using a 1,200-kN UTM at ramp rate of 0.6

mm/min as shown in Fig. 4 [27]. Five replicate

beams were tested for the BFRP- and the PEN FRP-

strengthened beams, respectively.

A small strip of BFRP or the PEN FRP was ad-

hered using adhesive to one end of all beam bond

specimens as shown in Fig. 3(c). After completion of

the beam bond tests, the BFRP or PEN FRP strip at-

tached at the end were subjected to pull-off test using

a pull-off testing device equipped with a 50 x 50 mm

square steel end plate. The steel end plate was ad-

hered to the BFRP or PEN FRP strip using two-part

epoxy after the FRPs were cut to fit 50 mm x 50 mm

steel end plate (See Fig. 5). Seven days after the ad-

hesive application, the pull-off test was performed as

shown in Fig. 5(d). Average bond stress in tension

was determined by dividing the maximum pull-off

load (Pmax) by the contact area (A = 2,500 mm2). The

pull-off test results of the FRPs retrieved from the

specimens immersed in water and 1N NaOH solution

and those stored in the room condition were com-

pared.

(a) Water (b) 1N NaOH (c) Room condition

Fig. 3 Beam bond test specimens immersed in water/alkaline solutions and stored in room condition


41

(a) A beam bond test specimen (b) Beam bond test in progress

Fig. 4 Beam bond test setup

(a) PEN FRP (b) BFRP (c) Steel end plate (d) Pull-off test

Fig. 5 Pull-off test

2.4 Tensile test of PEN fiber/FRP before/after

exposure to NaOH/NaCl solutions and wa-

ter

In this phase of study, durability characteristics

of PEN fiber/FRP (i.e. Uniaxial PEN fiber sheet and

PEN FRP) were investigated by observing weight

change and change in the tensile properties before

and after exposure to different environmental condi-

tions up to 6 months. Multiple PEN fiber/FRP tensile

specimens were prepared in length of 400 mm (PEN

fiber sheet) or 300 mm (PEN FRP). Test specimens

were immersed in three different environments: i.e.

Water (i.e. 100% R.H. at 40°C) and 3% NaCl and 1N

NaOH solutions at 20°C, respectively. The above en-

vironmental conditions simulated completely wet

condition, seawater, and sound concrete with high al-

kalinity, respectively.

The weight measurement and the tensile test

were carried out before the durability test began and

after immersion for 1, 3, and 6 months to determine

the weight loss and the strength/stiffness loss in ten-

sion, if any. After each planned immersion period,

test specimens were taken out of the environmental

chamber and dried for two days in a container about

quarter full of silica gel. After the specimens were

dried, they were weighed using a high precision scale

(accuracy = 0.0001 g) and then the tensile test was

performed following ISO 10406-2 (ten tests each).

Scanning electron microscopy (SEM) photos were

taken before/after the planned immersion period.

Figure 6 shows PEN fiber sheet and PEN FRP spec-

imens immersed in NaOH, NaCl solutions, or water.

(a) PEN fiber sheet (b) PEN FRP

Fig. 6 PEN fiber/FRP specimens immersed in NaOH/NaCl solutions or water


42

2.5 PEN fiber/FRP immersion in diluted acidic

solution

This specific durability test by tensile testing af-

ter immersion in acidic solution was carried out only

for PEN fiber/FRP. Tensile test coupons consisted of

PEN fiber sheet and PEN FRP, which were im-

mersed in diluted acidic solutions of 5% HCl or 10%

HCl for 3, 7, and 14 days. After the planned immer-

sion period, the specimens were retrieved, dried for

2 days and subjected to tensile test. Tensile strength

after immersion in diluted acidic solution was com-

pared to that of the control specimen. Three replicate

specimens were tested.

2.6 Exposure in natural environment for PEN

fiber/FRP

This test was carried out for PEN fiber/FRP

only. PEN fiber sheet and PEN FRP tensile test spec-

imens were prepared which were exposed in the out-

door conditions in South Korea starting from late

winter (T = 0°C ~ -10°C, R.H. = 30%-40% typ.) to

early summer (T = 20°C-30°C, R.H. = 40%-60%

typ.). After six-months’ exposure in the natural envi-

ronment including direct sun ray and U.V., the spec-

imens were retrieved and subjected to tensile test.

Change of tensile strength before/after exposure was

examined using ten replicate specimens of PEN fi-

ber/FRP, respectively.

3. Test results

3.1 Double lap shear test results

The double lap shear test was performed using

two different FRP composites adhered to concrete:

BFRP (2 layers) and PEN FRP (1 layer). Figure 7

schematically shows the location of the strain gauges

attached on the surface of FRP and the force acting

on FRP at center and in between the strain gauge lo-

cations. As soon as a test specimen shown in Fig. 2

is subjected to tensile force P, the concrete is sepa-

rated into two parts due to a thin steel plate placed at

center and, as a result, each FRP sheet bonded on the

side face of the concrete block is subject to a tensile

force of P/2 at center (loading end). Bond stress (u)

acting at FRP-concrete interface between ith and

(i+1)th strain gauges in Fig. 7 can be determined from

Eqs. (1) through (4), where the stress is determined

from strain readings of the strain gauge:

𝑇𝑖 = 𝜎𝑥𝑖(𝑡𝑦). ……………… (1)

𝑇𝑖+1 = 𝜎𝑥𝑖+1(𝑡𝑦)…………..… (2)

∆𝑇 = 𝑇𝑖 − 𝑇𝑖+1……………….(3)

𝑢 =∆𝑇

𝑥𝑦……………………...(4)

Where Ti = tensile force acting on FRP at ith

strain gauge, σxi = axial stress of FRP at ith strain

gauge location, t = thickness of FRP, x = distance

between two adjacent strain gauges, y = width of

FRP.

Table 3 summarizes the maximum axial strain

observed by the measured strain (εx-frp), correspond-

ing axial stress (σx-frp), tensile force acting on the FRP

layer (Pfrp), and the maximum bond stress (umax) de-

termined by Eqs. (1) through (4). Figures 8 and 9

show distribution of axial strains and bond stresses

for 126-mm distance starting from center for BFRP

and PEN FRP, respectively. It is observed from

Table 3 and Fig. 8 that the maximum axial strain,

axial stress, in Table 3 and Fig. 9, the maximum axial

strain, axial stress, FRP tensile force, and bond stress

are 0.7%, 149 MPa, 15.1 kN, and 3.0 MPa, respecti-

vely, for PEN FRP-1 while they are 0.82%, 172

MPa, 17.4 kN, and 4.3 MPa for PEN FRP-2. The dis-

tance the bond stress develops (Le) is 63-84 mm for

BFRP while it is 105 mm for PEN FRP. The distance

the bond stress develops as shown in Figs. 8 and 9

and Table 3 is known as the effective bond length Le

in literature: i.e. even if the bond length of the FRP

is larger than Le, the bond stress still develops within

Le. It can be seen that the effective bond length of

BFRP (63-84 mm) determined in this study agrees

well with that suggested by Shen et al., where the ef-

fective bond length varied between 56-72mm depen-

ding on the strain rate [8]. The bond stress between

BFRP and concrete at the interface is relatively high

and narrowly distributed as shown in Fig. 8 and

Table 3. On the other hand, the bond stress is rather

low and distributed over a larger distance for PEN

FRP in Fig. 9 and Table 3. In all tests, the failure

mode was debonding at the FRP-concrete interface

as shown in Fig. 10 and Table 3.

Fig. 7 Force (P/2) acting at center and away from center [21]


43

Table 3 Summary of test results at maximum fiber strain (εx-frp)

Fiber type εx-frp

σx-frp

(MPa)

Afrp

(mm2)

Pfrp

(kN)

Le

(mm)

umax

(MPa)

umean

(MPa)

umax /

umean

Failure mode

BFRP-1 0.0141 963 22.6 21.8 84 7.4 2.16 3.4 debonding

BFRP-2 0.0119 812 22.6 18.4 63 8.4 3.50 2.4 debonding

PEN FRP-1 0.0070 149 84.3 15.1 105 3.0 1.20 2.5 debonding

PEN FRP-2 0.0082 172 84.3 17.4 105 4.3 1.38 3.1 debonding

(a) BFRP-1 (a) BFRP-2

(b) BFRP-1 (b) BFRP-2

(c) BFRP-1 (c) BFRP-2

Fig. 8 Distribution of axial strain, axial stress, and bond stress at varying level of BFRP tensile force


44

(a) PEN FRP-1 (a) PEN FRP-2

(b) PEN FRP-1 (b) PEN FRP-2

(c) PEN FRP-1 (c) PEN FRP-2

Fig. 9 Distribution of axial strain, axial stress, and bond stress at varying level of PEN FRP tensile force

3.2 Beam-bond test results of BFRP and PEN

FRP

The beam bond test specimens were retrieved

after 4-months’ immersion in water or alkaline solu-

tion (1N NaOH) as well as the control specimens

(stored in room condition for 4 months) and tested

by flexure test by three-point loading following ACI

440.9R [27]. As the applied load increased, a tensile

crack quickly appeared and developed upward star-

ting from top of the notch (See Fig. 4). At peak load,

BFRP- and PEN FRP-strengthened beams failed ei-

ther by debonding at the FRP-concrete interface or

FRP fracture in tension. Average peak load of C-

BFRP (Control), RH-BFRP (R.H. 100%), and

NaOH-BFP (1N NaOH) was 15.5 kN, 15.7 kN, and

7.94 kN, respectively, as shown in Table 4. Average

peak load for C-PEN FRP, RH-PEN FRP, and


45

NaOH-PEN FRP was 19.7 kN, 17.4 kN, and 17.4

kN, respectively. Equation (5) was used to determine

beam bond retention (or residual mechanical pro-

perty) as suggested by ACI 440.9R [27].

𝑅𝑒𝑏 =𝑃𝑏2

𝑃𝑏1× 100 (%)………… (5)

Where, Pb2 is peak load of water/alkali condi-

tioned specimen, Pb1 is peak load of Control speci-

men and Reb is beam is beam bond retention.

Table 4 Results of beam bond and pull-off tests

Index Beam bond test Pull-off

Test

(MPa)

Index Beam bond test Pull-off

Test

(MPa) Load

(kN)

Reb

(%)

Load

(kN)

Reb

(%)

C-BFRP 15.5 -- 5.23 C-PEN FRP 19.7 -- 4.47

RH-BFRP 15.7 100 3.97 RH-PEN FRP 17.4 88 3.56

NaOH-BFRP 7.94 51.2 3.55 NaOH-PEN FRP 17.4 88 1.90

NOTE: Results shown are average of 5 tests.

(a) PEN FRP (b) BFRP (c) Debonding failure

Fig. 10 Double lap shear test in progress and failure mode

Table 4 summarizes the beam bond test results in

terms of peak load and beam bond retention. The

beam bond retention is 88% for PEN FRP both for

wet condition and in alkaline condition. On the other

hand, the beam bond retention is 100% and 51% in

wet condition and in alkaline condition, respectively,

for BFRP. Test results indicate that BFRP may be

vulnerable for alkaline exposure. Figures 11 and 12

show the beam bond test specimens after test. In Fig-

ure 11 which shows the control specimens, the fail-

ure occurred at the FRP-concrete interface both for

BFRP and PEN FRP. C-PEN FRP shows some con-

crete attached to the debonded FRP surface while C-

BFRP shows almost complete and clean interface

failure. Figure 12 shows top surfaces of C-BFRP,

RH-BFRP, and NaOH-BFRP after beam bond test.

C-BFRP retains original dark brown color of the

BFRP in Fig. 12(a). On the other hand, yellowish

discoloration of BFRP is shown for RH-BFRP in

Fig. 12(b) while the discoloration is more severe for

NaOH-BFRP in Fig. 12(c). It has been reported that

the basalt fiber can be degraded under soaked, saline,

and alkaline environment while the degradation of

the basalt fiber can be significantly slowed down by

oxide coating such as zirconium deoxide and tita-

nium deoxide coating [12-14]. Since the basalt fibers

used in this study were not coated, further investiga-

tion on the durability properties of the basalt fiber

was not pursued.

As described in Clause 2.3, the pull-off test of

BFRP or PEN FRP bonded to the end of the beam

bond test specimens was also performed after the

completion of the beam bond test. Table 4 and Fig-

ures 13 and 14 show the results of pull-off test. Pull-

off test results indicate that both immersion in water

and alkaline solution at an elevated temperature tend

to degrade the bond between concrete and

BFRP/PEN FRP. C-BFRP (Control), RH-BFRP, and

NaOH-BFRP have average pull-off bond strength of

5.23 MPa, 3.97 MPa, and 3.55 MPa, respectively, in

Table 4, while average bond strength of C-PEN FRP

(Control), RH-PEN FRP, and NaOH-PEN FRP is

4.47 MPa, 3.56 MPa, and 1.90 MPa, respectively. C-

BFRP specimens failed mostly in concrete substrate

while NaOH-BFRP specimens failed partially in

concrete substrate. RH-BFRP specimens failed

mostly at interface. PEN FRP pull-off specimens

show partial failure at interface and concrete sub-

strate for C-PEN FRP, and clean interface failure for


46

RH-PEN FRP and NaOH-PEN FRP specimens indi-

cating damage at the interface by 4-months’ immer-

sion in water and alkaline solutions.

3.3 Test results of PEN fiber/FRP before/after

exposure to NaOH, NaCl solutions and wa-

ter

Table 5 compares the weights/weight losses be-

fore and after 30, 90, and 180 days of exposure for

PEN fiber sheet and PEN FRP under three different

conditions of completely wet (Wet, 40°C) and 3%

NaCl and 1N NaOH solutions (20°C), respectively.

In Table 5, PEN fiber sheet shows 0.33%, 0.18%,

and 10.5% weight loss after exposure to Wet, NaCl,

and NaOH conditions, respectively, while the loss is

0.80%, 0.76%, and 4.34% for PEN FRP, respec-

tively, after 180 days. PEN fiber shows some nega-

tive effect under alkaline environment. However, the

possible negative effect of alkaline environment on

PEN fiber is significantly reduced in case of PEN

FRP as the PEN fiber is embedded in the adhesive

matrix.

(a) C-BFRP (b) C-PEN FRP

Fig. 11 Control specimens after beam bond test

(a) C-BFRP (b) RH-BFRP (c) NaOH-BFRP

Fig. 12 BFRP specimens after beam bond test

Table 6 summarizes the tensile test results for

PEN fiber sheet and PEN FRP in terms of tensile

strength and elastic modulus before and after 30, 90,

and 180 days. In Table 6, the tensile strength of PEN

fiber sheet after 180 days is 114%, 115%, and 97%

of control when exposed to Wet, 3% NaCl, and 1N

NaOH conditions, respectively. The tensile strength

for PEN FRP after 180 days is 98%, 96%, 93% for

Wet, 3% NaCl, and 1N NaOH conditions, respec-

tively. Elastic modulus is also shown in Table 6. The

elastic modulus (E1) after 180 days ranges 98%-

103% for PEN fiber sheet and 100%-104% for PEN

FRP, respectively. It can be concluded that the PEN

fiber/FRP has satisfactory resistance after exposure

to all three environmental conditions of Wet, 3%

NaCl, and 1N NaOH after 180 days.

Figure 15 shows SEM images before and after

180 days of exposure to different environmental con-

ditions, where no significant change is visible be-

fore/after exposure. In Figure 15(a), the longitudinal

surface of PEN fiber filament is very smooth and

clean while the diameter is 30 μm before exposure.

In Figure 15(b)-(d), after 180 days’ exposure to wet,

saline, and alkaline conditions, the surfaces do not

show any significant trace of chemical etching or

bruises in all conditions.


47

3.4 Influence of acidic condition

Figure 16 compares the test results of PEN fi-

ber/FRP before and after immersion in 5% and 10%

diluted solutions of HCl for 3, 7, and 14 days: i.e.

Tensile strengths after immersion are normalized in

terms of tensile strength before immersion in Fig. 16.

It is seen in Fig. 16(a) that the tensile strength of PEN

fiber sheet after immersion tends to be affected a lit-

tle, while the tensile strengths are almost unaffected

in Fig. 16(b) for PEN FRP specimens.

3.5 Exposure to natural environment

As described earlier, PEN fiber sheet and PEN

FRP tensile test specimens were prepared and ex-

posed in natural outdoor conditions for six months.

Tensile test results are summarized in Table 7 after

six months’ exposure in the natural environment in-

cluding direct sun ray and U.V. Degradation is

clearly noticed for PEN fiber sheet while the degra-

dation is significantly reduced for PEN FRP in Table

7.

Fig. 13 Pull-off test results

(a) C-BFRP (b) RH-BFRP (c) NaOH-BFRP

(d) C-PEN FRP (e) RH-PEN FRP (f) NaOH-PEN FRP

Fig. 14 Failure mode of pull-off test specimens


48

Table 5 Summary of weight measured before/after exposure to wet, 3% NaCl and 1N NaOH conditions

Fiber type Environ-

ment

Temp.

(̊C)

Weight

before

(g)

Weight af-

ter 180d

(g)

Weight loss after (%)

30d 90d 180d

PEN fiber

sheet

Wet

NaCl

NaOH

40

20

20

7.7040

7.6722

7.6815

7.6784

7.6583

6.8789

0.37

0.02

1.64

0.22

0.02

5.51

0.33

0.18

10.5

PEN FRP

Wet

NaCl

NaOH

40

20

20

11.177

11.007

11.083

11.088

10.924

10.602

0.44

0.16

0.73

--

0.05

1.70

0.80

0.76

4.34

NOTE: PEN uniaxial fiber sheet consists of 6 PEN fiber rovings; 2. Specimen length is 400 mm for PEN fi-

ber roving, and PEN fiber sheet and 300 mm for PEN FRP; 3. Adhesive amount is 150% of fiber by vol. for

PEN FRP; 4. Average of twelve measurements

Table 6 Summary of tensile test results before/after exposure to wet, 3% NaCl and 1N NaOH conditions

Fiber

type

Environ-

ment

Temp.

(°C)

Tensile strength before/after

(MPa)

Elastic modulus before/after

(GPa)

before 30d 90d 180d before 30d 90d 180d

PEN fi-

ber sheet

Wet

NaCl

NaOH

40

20

20

842

786

923

861

932

935

851

961

966

816

17.5

15.9

17.2

17.0

17.3

16.9

17.7

18.0

17.9

17.2

PEN FRP

Wet

NaCl

NaOH

40

20

20

912

893

952

857

902

935

900

897

880

844

21.4

20.7

20.3

21.9

19.7

20.0

20.3

21.3

21.4

22.2

NOTE: Average of twelve tests

(a) Control (b) Wet

(c) 3% NaCl (d) 1N NaOH

Fig. 15 SEM photos of PEN after exposure for180 days to different environmental conditions


49

(a) PEN fiber sheet (b) PEN FRP

Fig. 16 Results of tensile test of PEN fiber/FRP after immersion in acidic solutions

Table 7 Summary of tensile strengths for PEN FRP after exposure in natural conditions for up to 6 months

Fiber type Before

(MPa)

After 6 months

(MPa)

After/Before

(%)

PEN fiber sheet 842 594 70.5

PEN FRP 912 814 89.3

NOTE: Average of 10 tests

4. Conclusions

Bond between BFRP-to-concrete and PEN

FRP-to-concrete interface was studied in terms of

distribution and magnitude of bond stress by double

lap shear test. In addition, durability characteristics

of BFRP and PEN fiber/FRP have been studied by

(1) beam bond test following accelerated condition-

ing protocols for durability assessment by ACI

440.9R, (2) fiber tensile test before/after immersion

in 1N NaOH and 3% NaCl solutions, and water, (3)

fiber tensile test before/after immersion in diluted

acidic solutions, and (4) exposure to natural outdoor

environment. The durability investigation by beam

bond test included both BFRP and PEN FRP. Other

durability investigations concentrated on PEN fi-

ber/FRP. The following conclusions can be drawn

from the current experimental study.

(1) Double lap shear test results show that the bond

stress that develops at the interface between

BFRP and concrete is relatively high and nar-

rowly distributed: Maximum bond stress was 8.4

MPa and distance the bond stress distributed was

63-84 mm for BFRP. On the other hand, the

bond stress was relatively low and distributed

over a larger distance for PEN FRP: Maximum

bond stress was 4.3 MPa while the distance the

bond stress distributed was 105 mm for PEN

FRP. Test results show that the effective bond

length Le is 84 mm for BFRP and 105 mm for

PEN FRP.

(2) Beam bond test results show that the behavior of

PEN FRP under wet and alkaline conditions is

relatively good with beam bond retention of 88%

after immersion in water and alkaline solutions

for 4 months. Behavior of BFRP was not as good

under alkaline environment with the beam bond

retention of 51.5% after immersion for 4 months.

(3) Results of weight measurement and tensile test

of PEN fiber/FRP before and after wet, alkaline,

and saline conditions indicated possible degrad-

ing of PEN fiber under alkaline condition. For

the PEN FRP, however, the weight reduction

(about 4%) and reduction of the tensile strength

was small (7.5%) and there was no change of the

elastic modulus before/after immersion for 6

months.

(4) Tensile strength of PEN FRP immersed in 5%

and 10% diluted acidic solutions of HCl was

very good after 14 days: no reduction in 5% so-

lution and less than 3% reduction in 10% solu-

tion. Tensile strength of PEN FRP after exposure

for 6 months in natural environment showed

about 10% reduction.

(5) In overall, the performance of PEN fiber/FRP

was satisfactory in all durability tests conducted

in this study.

Acknowledgements

This study was supported by Grant from Korea

Re Research Foundation, Grant number [NRF2018-

R1D1A1B07049635]. Authors gratefully acknowle-

dge the generous support


50

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Brijesh Singh, Vikas Patel*, P N Ojha, VV Arora · 2020-07-16 · Brijesh Singh, Vikas Patel*, P N Ojha, VV Arora ... [1,2]. However, high-strength concrete is more brittle in nature

Brijesh Singh, Vikas Patel, P N Ojha, VV Arora · 2020-07-16 · Brijesh Singh, Vikas Patel, P N Ojha, VV Arora ... [1,2]. However, high-strength concrete is more brittle in nature