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BOND BEHAVIOR OF REINFORCING BARS ON LOW BINDER
AND RECYCLED AGGREGATE CONCRETE
Tiago Amândio Santos Pereira
Instituto Superior Técnico, May 2019
ABSTRACT
The CO2 emissions from the production of cement and the large-scale production of construction and
demolition waste (CDW) have led the scientific community to seek alternative solutions in order to
achieve the environmental and economic sustainability of the construction sector. At the structural level,
two of the possible proposed solutions are low binder concrete (LBC) and low cement and recycled
aggregate concrete (LCRAC).
This dissertation intends to analyze the viability of the structural use of LBC and LCRAC by studying the
bond behavior of each of the concrete’s to reinforcing bars. To that end, the experimental program
includes the performance of monotonic pull-out tests, testing the influence of LCB and LCRAC mixtures
on bond behavior, and, in addition, the influence of packing density {0,82; 0,84; 0,86}, the aggregates’
optimization distribution (Alfred and Faury curve) and the incorporation of recycled aggregates {30%;
55%; 80%}.
From these experimental procedures, the following main conclusions were made: (1) LBC and LCRAC’s
packing densities have a positive effect on bond behavior at serviceability and ultimate limit states, and
(2) whereas natural aggregates’ size should be smaller than the rebar’s ribs spacing, the size of recycled
aggregates must be larger than such seeing as, otherwise, recycled aggregates may be harmful to the
bond mechanism between concrete and rebars.
Furthermore, this study asserts, based on four unique studies and in the Model Code 2010, a new and
innovative proposal for the bond-slip relationship law that contemplates every variable under discussion.
Key-words: sustainability; low binder/cement; packing density; recycled aggregate of CDW; bond
behavior rebar-concrete; bond-slip relationship
2
1. INTRODUCTION
Concrete, whose most essential component is
cement, is vital to the engineering and construction
sector due to its competitive market price and
structural behavior. It is the second most widely
used material in the world after water, having an
annual production of approximately 4.65 billion tons
with an anticipated growth of 12 to 23% by 2050.
Today, however, the production of Portland cement
is associated with environmentally detrimental
emissions. For every ton of Portland cement
produced, approximately one ton of CO2 is emitted,
accounting for about 5% of global greenhouse gas
emissions [1].
Growing environmental awareness has encouraged
the rise in strategies aimed at reducing pollutant
emissions from cement production, integrating
sustainable solutions, concerning the cement
manufacturing process, and sustainable
substitutes, or additions. Recent studies on low
binder concrete (LBC), which partially decrease and
replace the amount of cement by additions from
recovered industrial by-product, reveal that LBC is
an excellent alternative to current concrete,
showcasing identical strength and even better
durability [2] [3].
The amount of construction and demolition waste
(CDW) generated by the construction sector
accounts for about 25 to 30% of the European
Union’s. As such, the European Union has been
imposing restrictions that encourage the recycling of
these wastes. However, one of the obstacles that
has yet to be overcome in the reuse of CDW is the
ambiguity associated with its quality. Its composition
varies according to the origin of the incorporated
residues, leading to insufficient knowledge of the
properties thereof. However, the efforts achieved by
the scientific community throughout this century
highlight CDW’s potential as an aggregate for
concrete, as its deficient use is mitigated despite
being subjected to certain methodologies [4] [5].
Regardless of all the established evidence for new
strategies designed to make concrete more
environmentally sustainable, the structural behaviour
of this “greener” concrete remains partially unknown,
specifically with regards to its adherence and how
this particular property is influenced by different
formulations of the concrete. Therefore, this paper
experimentally investigates concrete’s bond
behavior, in particular how its influenced by: packing
density {0.82; 0.84; 0.86}, aggregate optimization
based on the Faury improved and the Alfred
reference curves, and the incorporation of recycled
aggregates {30%; 55%; 80}.
In addition, this paper proposes a set of expressions
that redefine the law of the bond-slip relationship,
advocated by the scientific community and current
European standards. These formulas integrate the
findings of variable studies and enable a more
realistic calculation of anchorage zones and plastic
spherical plain bearings.
2. EXPERIMENTAL CAMPAIGN
The influences of packing density, the particle size
optimization curve, and the substitution of NA by RA
on the bond between rebar and LBC/LCRAC were
tested.
Concrete mixtures
LBC and LCRAC had low dosages of binder and
cement, respectably. The dosages used were less
than 260 kg/m3, the recommended minimum by NP
EN 206-1 [6].
The constituent aggregates utilized include fine silty
sand of 0/1 mm, average sand of 0/4 mm, rolled
river sand of 4/8 mm, limestone gravel of 6/14 mm,
and three distinct fractions of CDW: 1/20 (2350
kg/m3), 4/20 (2350 kg/m3) and 10/20 mm (2220
kg/m3).
3
Through the optimization of the granulometric curve
of the aggregates different packing densities were
attained. Particularly, this was allowed by Alfred's
reference curve and the corrected Faury’s curve.
The corrected Faury’s curve uses larger amounts of
coarse aggregates than the original Alfred’s curve.
This higher incidence of coarse aggregates has a
tendency to increase voids between aggregates,
voids which will be compensated for with a greater
amount of fine sand, which in turn increases binder
demands. On the other hand, Alves [2] used the
corrected Faury granulometric curve without
compromising the amount of binder.
Through Alfred´s curve, three packing densities
(0.82, 0.84, and 0.86) were employed in LBC. In
addition, an additional packing density of 0.86 was
achieved with the corrected version of Faury’s
curve.
Fresh and hardened concrete procedures
The aggregates, with or without RA, were the first to
be placed in the concrete mixer, undergoing a dry
mix for 45 seconds. Subsequently, 1/4 of the water
content was added to the LBC. Conversely, the
LCRAC received half of the water content and was
mixed together with the aggregates for 2 minutes,
allowing for NA’s partial absorption of water and the
resulting correct hydration of the cement (water
compensation method). The remaining water,
diluted with the superplasticizer in it, was
simultaneously added to both cements, ensuring
the required workability.
Alves [2], Freitas [3], Carvalho [7], and Robalo [8]
characterized the fresh and hardened states of LBC
and LCRAC and the durability of LBC with
compositions similar to those studied in this paper.
Once the LBC and the LCRAC were characterized,
the test specimens produced for the
characterization of concrete were only aimed at
confirming the quality of the concrete mixture and at
learning the compressive strength, at the concrete’s
age, through the pull-out test.
For this purpose, 24 (three for each LBC and
LCRAC) cubic specimens of 150x150x150 mm and
8 (three for each LBC) cylindrical specimens (Φ =
150 mm and H = 300 mm) were produced for the
compressive strength test and for the diametral
compressive strength test, respectively.
All compressive strength and concrete tests
followed the methodology prescribed in NP EN
12390-3 [9]. All tensile strength diametral
compressive tests followed the methodology
prescribed in NP EN 12390-6 [10].
The concrete strength at 28 days, fcm,28, was
predicted through the Euro code 2 [11] hardening
curve and is given by Eq. (1):
𝑓𝑐𝑚,𝑗(𝑗) = 𝑒{𝑠∙[1−(
28𝑗
)1/2
]}∙ 𝑓𝑐𝑚,28
(1)
where s has a recommended value of 0.60 for LBC
and 0.25 for LCRAC by Alves [2] and Robalo [8],
respectively.
Rebars
All rebars were subjected to a rib measurement test,
according to section 10 of EN ISO 15630-1 [12] and
the LNEC E460 specification [13]. The
measurements were made in each specimen bond
length, in order to reduce experimental error due to
the ribs’ influence on adhesion.
The relative rib area, fR, was calculated, using the
experimental results, by Eq. (2):
𝑓𝑅 =2 ∙ 𝑎1/2
3 ∙ 𝜋 ∙ Φ ∙ 𝑐∙ (𝜋 ∙ Φ − ∑ 𝑓𝑖) (2)
where, Φ is the bar diameter, c is the rib spacing,
a1/2 is the ribs’ height, and ∑ 𝑓𝑖 is the part of the
circumference without rib.
4
Bond behavior
35 pull-out tests (7 different concretes, 5 for each)
of 200x200x200 mm specimens, with partially
embedded rebars, were carried out. All embedment
lengths were 5Φ mm.
The tests were assisted by a servo hydraulic jack
(Schenck), an accessory for the installation of test
specimens, a measurement system integrated by a
linear displacement transducer (HWB WA20), a
measurement transducer fastened to the specimen,
and an automatic data acquisition unit (Spider 8
incorporated with software Catman32) (Figure 1).
Figure 1 – Pull-out test system.
The methodology followed the recommended in
annex D of EN 10080 [14], which is based on RILEM
– Recommendation RC 6 Bond test for
reinforcement steel – 2. Pull-Out Test [15], except
for the load application rate and the calculation of
bond strength. All pull-out tests were performed at
the rate of 0.03 mm/s, as suggested by Eligehausen
[16], leading to plausible and conservative results.
The bond strengths were given by the Model Code
2010, Eq. (3):
𝜏𝑑 =𝐹𝑎
5 ∙ 𝜋 ∙ Φ2∙ √
𝑓𝑐𝑚,28
𝑓𝑐𝑚,𝑗
(3)
where, Fa is the testing force, and fcm,28 and fjcm,j are
the concrete strength at 28 days and at the testing
day, respectively.
Annex D of EN 10080 [14] presents the calculation
of bond strength with 𝑓𝑐𝑚,28 𝑓𝑐𝑚,𝑗⁄ instead of
√𝑓𝑐𝑚,28 𝑓𝑐𝑚,𝑗⁄ ; however, the Model Code 2010 [17]
considers the non-linear dependence of
compressive strength on bond behavior.
The average bond strength, τd,average, is calculated
according to Annex C of the Eurocode 2 [11], by the
following expression (4),
𝜏𝑑,𝑚é𝑑𝑖𝑎 =𝜏𝑑,0,01 + 𝜏𝑑,0,1 + 𝜏𝑑,1,0
3 (4)
where, τd,0,01, τd,0,1, and τd,1,0 represent the bond
strength at 0.01 mm, 0.1 mm and 1.0 mm of slip,
respectively.
3. RESULTS AND DISCUSSION
Fresh and hardened concrete
Each concrete’s average compressive strengths at
day 28 and at the pull-out testing day j, and average
tensile strengths are shown in Table 1.
Rebars
The average values for the rebar’s ribs geometric
characteristics were compared with the values
specified for their certification, according to the
LNEC specification E460 (2010) and annex C of the
National Annex (NA) and the European Standard
(EN) of the Eurocode 2 (EC2) [11] (Table 2).
5
Table 1 – Average concrete compressive strengths at day 28 and at pull-out testing day j and average concrete tensile
strengths.
j
[days] Density at j days
[kg/m3] fcm,cube,j [MPa]
fcm,j [MPa]
fcm,28 [MPa]
fctm,28 [MPa]
𝒇𝒄𝒕𝒎,𝒋
𝒇𝒄𝒎,𝒄𝒖𝒃𝒐,𝒋 [%]
LBC_0,86_Alfred 28 2394 36,2 29,7 29,7 2,6 7,3
LBC_0,84_Alfred 29 2366 24,8 20,4 20,2 2,2 9,1
LBC_0,82_Alfred 30 2344 18,3 15,0 14,7 1,8 9,9
LBC_0,86_Faury 33 2422 34,3 28,2 26,9 2,4 7,0
LCRAC_30 28 2276 30,6 24,5 24,5 - -
LCRAC_55 31 2207 24,5 19,6 19,3 - -
LCRAC_80 32 2144 19,7 15,8 15,5 - -
(*) 𝑓𝑐𝑚 = 𝑘1 ∙ 𝑓𝑐𝑚,𝑐𝑢𝑏𝑒, where k1 is 0,82 for LBC [18] and 0,80 for LCRAC [6].
Table 2 – Geometric characteristics of the rebar’s ribs.
Parameter
Specified value
Mean Standard deviation
Minimun Maximum E460
EC2
NA EN
a1/2 [mm] ≥ 0,78 ≥ 0,78 - 0,95 0,013 0,92 0,97
c [mm] 7,2 ± 15% 7,2 ± 15% - 7,7 0,06 7,4 7,8
Σfi [mm] ≤ 7,5 ≤ 7,5 - 4,2 0,27 3,7 4,9
fR ≥ 0,056 ≥ 0,056 ≥ 0,040 0,073 0,0015 0,070 0,077
a1/2 [mm] ≥ 0,78 ≥ 0,78 - 0,93 - 0,92 0,93
c [mm] 7,2 ± 15% 7,2 ± 15% - 7,7 - 7,7 7,8
Σfi [mm] ≤ 7,5 ≤ 7,5 - 4,2 - 4,2 4,3
fR ≥ 0,056 ≥ 0,056 ≥ 0,040 0,071 - 0,070 0,072
Bond behavior on low binder
All specimens’ failure mode was pull-out failure. In
Figure 2, the bond strength-slip relationship of the
LBC_0.86_Alfred is illustrated.
Figure 2 – Bond-slip relationship of LBC_0,86_Alfred.
Through the analysis of the packing density and the
optimization aggregate curve, the fcm,28 were
standardized at 25 MPa (Eq. 3) isolating the
variables, independently of the concrete’s
compressive strength.
3.3.1 Influence of the optimization
aggregate curve
The superiority in bond strength of the
LBC_0.86_Alfred over LBC_0.86_Faury was
attributed to the interlocking between the rebar ribs
and the aggregates accommodated between them.
As a result, the quantity of aggregates with
interlocking potential, i.e., smaller in size than the rib
spacing, c, equal to 7.7 mm, was evaluated.
Considering the results of this experimental
campaign, the addition of 9% of aggregates, with a
smaller size than c, reflected an increase of 7%⬆
and 6%⬆ of the τd,average and the τd,max, respectively.
Notwithstanding the slight differences between LBC
mixtures, Freitas’ [3] results were also considered.
An additional 14% of aggregates smaller in size
than c resulted in an increase of 12%⬆ in τd,max.
However, the interlocking hypothesis between the
0
5
10
15
20
25
30
0 2 4 6 8 10 12 14 16 18 20
Bo
nd
str
en
gth
[M
Pa
]
Slip [mm]
fcm,28 = 29,7 MPa
1º POT
3º POT
4º POT
5º POT
6
aggregates and the ribs was not observed in the
τd,average from Freitas’ results [3].
3.3.2 Influence of packing density
The bond strength-packing density relationship of
LBC_Alfred is shown in Figure 3. The τd,average and
τd,max increased directly with packing density. The
LBC_0,86_Alfred indicated an increase of 79%⬆
and 37%⬆, in relation to LBC_0,82_Alfred.
The relationship between bond strength, either
τd,average or τd,máx, and packing density was linear.
However, LBC_0,84_Alfred’s τd,max value brought
into question the possibility for a future study on the
perhaps not so linear nature of the relationship,
similar to that of concrete’s compressive strength
dependence on bond behavior.
Notwithstanding the diversity of LBC mixtures,
Freitas’ results [3] were also regarded and are
shown in Figure 4. In such, the addition of 0.05 to
the packing density increased the τd,max by 38%⬆.
On the other hand, the τd,average revealed two trends:
(i) a more expressive one largely based on this
dissertation’s results, where the additional 0.05 in
packing density increased the τd,max by 98%⬆; (ii)
and a more modest one largely based on Freitas’
results [3], where the τd,max increased by 28%⬆.
Figure 3 – Average (left) and maximum (right) bond strength-packing density relationship.
Figura 4 – Average (left) and maximum (right) bond strength-packing density relationship according to Freitas’ results [3].
τd,média = 184,4∙σ - 141,70
R² = 1,00
5,0
10,0
15,0
20,0
0,81 0,82 0,83 0,84 0,85 0,86 0,87
Ave
rag
e b
on
d s
tre
ng
th [M
Pa
]
Packing density
LBC_0,86_Alfred
LBC_0,84_Alfred
LBC_0,82_Alfred
τd,máx = 163,2∙σ - 115,78
R² = 0,93
15,0
20,0
25,0
30,0
0,81 0,82 0,83 0,84 0,85 0,86 0,87
Ma
xim
imu
m b
on
d s
tre
ng
th [M
Pa
]
Packing density
LBC_0,86_Alfred
LBC_0,84_Alfred
LBC_0,82_Alfred
τd,média = 160,5∙σ - 121,78
R² = 0,98
5,0
10,0
15,0
20,0
25,0
0,80 0,81 0,82 0,83 0,84 0,85 0,86 0,87
Ave
rag
e b
on
d s
tre
ng
th [M
Pa
]
Packing density
C250LBC125LBC75LBC_0,86_AlfredLBC_0,84_AlfredLBC_0,82_AlfredLBC_0,86_Faury
τd,máx = 134,9∙σ - 91,54
R² = 0,79
15,0
20,0
25,0
30,0
35,0
0,80 0,81 0,82 0,83 0,84 0,85 0,86 0,87
Ma
xim
um
bo
nd
str
en
gth
[M
Pa
]
Packing density
C250LBC125LBC75LBC_0,86_AlfredLBC_0,84_AlfredLBC_0,82_AlfredLBC_0,86_Faury
fcm,28 = 25 MPa fcm,28 = 25 MPa
fcm,28 = 25 MPa fcm,28 = 25 MPa
7
Bond behavior on low cement and recycled aggregate concrete
All specimens’ failure mode was pull-out failure. In
Figure 5, the bond strength-slip relationship of the
LCRAC_30 is portrayed.
Figure 5 – Bond-slip relationship of LCRAC_30.
Through the analysis of the incorporation of RA,
fcm,28 were standardized at 25 MPa (Eq. 3) to isolate
the variables, independently of the concrete’s
compressive strength.
3.4.1 Influence of RA
Bond strength decreased with the continuous
incorporation of RA, contradicting previous studies
[19-24]. Hence, the possibility of the bond strengths
decreasing abruptly due to the incorporation of RA
smaller in size than the rib spacings arose.
The bond strength-incorporation of RA smaller in
size than the ribs spacings relationship is shown in
Figure 6. The incorporation of 34% of RA with
smaller sizes than the rib spacings showcased a
reduction in τd,average and τd,max of 25%⬇ and 32%⬇,
respectively.
Bond behavior prediction
Based on four unique studies [3] [25] [26] and in the
Model Code 2010 [17], a new innovative proposal
for the bond-slip relationship law, which
contemplates the influences of packing density,
relative rib area, and recycled aggregates
incorporation, was established.
This law is valid for "good" bond conditions and pull-
out failure modes and is conservative for rebar
diameters inferior to Φ25. It should be noted that, in
the presence of high compacities and transverse
ribs with relatively high areas and with 0%
incorporated RA smaller in size than ribs spacings,
maximum bond strengths, calculated with the
proposed law in equation (5), are abruptly 85%⬆
higher than those calculated by the Model Code
2010 [17].
Figure 6 – Average (left) and maximum (right) bond strength- incorporation of RA with smaller sizes than the ribs
spacings relationship.
0
5
10
15
20
25
30
0 2 4 6 8 10 12 14 16 18 20
Bo
nd
str
en
gth
[M
Pa
]
Slip [mm]
fcm,28 = 24,5 MPa
21ºPOT
22ºPOT
23ºPOT
24ºPOT
25ºPOT
τd,média = -10,9∙x + 15,28 (*)
R² = 0,92
5,0
10,0
15,0
20,0
0 20 40 60 80 100
Te
nsã
o m
éd
ia d
e a
de
rên
cia
[M
Pa
]
Σ RA < c [%]
LCRAC_30LCRAC_55LCRAC_80
τd,máx = -19,2∙x + 18,63 (*)
R² = 0,90
10,0
15,0
20,0
25,0
0 20 40 60 80 100
Te
nsã
o m
áxm
ia d
e a
de
rên
cia
[M
Pa
]
Σ RA < c [%]
LCRAC_30LCRAC_55LCRAC_80
fcm,28 = 25 MPa fcm,28 = 25 MPa
8
Table 3 – New proposal of the bond-slip relationship law.
Parameter New proposal of the bond-slip relationship law
τd,máx [MPa] 2,5 ∙ 𝜂𝑓𝑅∙ 𝜂𝜎 ∙ 𝜂𝑅𝐴<𝑐 ∙ √𝑓𝑐𝑚,28 (5)
ηfR (4757 ∙ 𝑓𝑅 + 878,5) ∙ 10−3 (6)
ησ (7927 ∙ 𝜎 − 5408) ∙ 10−3 (7)
ηRA<c (1000 − 972 ∙ 𝑅𝐴<𝑐) ∙ 10−3 (8)
s1 [mm] 1,0 − 5 ∙ (𝜎 − 0,82) if 𝜎 ≥ 0,82 (9)
s2 [mm] 2,0 − 5 ∙ (𝜎 − 0,82) if 𝜎 ≥ 0,82 (10)
s3 [mm] c (*)
α 0,4 − 4 ∙ (𝜎 − 0,82) if 𝜎 ≥ 0,82 (11)
τd,atrito [MPa] 0,30 ∙ 𝜏𝑑,𝑚á𝑥 for low cement concrete (12)
0,40 ∙ 𝜏𝑑,𝑚á𝑥 for reference concrete (13)
4. CONCLUSIONS
This study has led to the main following
conclusions:
(1) The LBC with high packing density was
characterized by the slight improvements in
workability compared to the homologous LBC
developed in previous studies [2] [3] [7].
(2) The LCRAC’s matrices demonstrated greater
sensitivity due to the increased water content
integrated for RA absorption. This became more
evident as more RA was incorporated.
(3) The use of the s coefficient, as recommended in
the Eurocode 2 [11] for the prediction of strength
at a certain age, was unviable.
(4) The influence of the LBC’s natural aggregates’
particle distribution on bond behavior appeared
to be slightly significant. However, the
incorporation of natural aggregates smaller in
size than the ribs spacing was determined
advisable.
(5) With regards to this study’s experimental
results, bond behavior underwent extreme
improvement with an increase in packing
density, i.e., the average bond strength, the
maximum bond strength, and the stiffness of the
ascending section of the bond-slip relationship
was significantly increased.
(6) With regards to Freitas’ experimental LBC [3]
results, both average bond strength and the
stiffness of the ascending section of the bond-
slip relationship demonstrated minimal
improvement compared to that observed on this
study’s experimental results.
(7) The influence of the incorporation of RA smaller
in size than the ribs spacing on bond behavior
appeared to be truly effective at reducing the
interlocking phenomenon’s influence between
concrete and aggregates. Therefore, the
incorporation of RA larger in size than the ribs
spacing is recommended.
(8) Based on four unique studies [3] [25] [26] and in
the Model Code 2010 [17], a new innovative
proposal for the bond-slip relationship law was
established, taking into consideration every
variable under discussion (Table 3).
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