1. ACI
MATERIALS
J O U R N A L
SPECIAL ISSUE:
ADVANCES IN RHEOLOGY AND
ADDITIVE MANUFACTURING IN CONSTRUCTION
V. 118, NO. 6, NOVEMBER 2021
@seismicisolation
@seismicisolation
3. 2 ACI Materials Journal/November 2021
Permission is granted by the American Concrete Institute for libraries and other users registered with the Copyright
Clearance Center (CCC) to photocopy any article contained herein for a fee of $3.00 per copy of the article. Payments
should be sent directly to the Copyright Clearance Center, 21 Congress Street, Salem, MA 01970. ISSN 0889-3241/98
$3.00. Copying done for other than personal or internal reference use without the express written permission of the
American Concrete Institute is prohibited. Requests for special permission or bulk copying should be addressed to the
Managing Editor, ACI Materials Journal, American Concrete Institute.
The Institute is not responsible for statements or opinions expressed in its publications. Institute publications are not able
to, nor intend to, supplant individual training, responsibility, or judgment of the user, or the supplier, of the information
presented.
Papers appearing in the ACI Materials Journal are reviewed according to the Institute’s Publication Policy by individuals
expert in the subject area of the papers.
Contributions to
ACI Materials Journal
The ACI Materials Journal is an open
forum on concrete technology and papers
related to this field are always welcome.
All material submitted for possible publi-
cation must meet the requirements of
the “American Concrete Institute Publi-
cation Policy” and “Author Guidelines
and Submission Procedures.” Prospective
authors should request a copy of the Policy
and Guidelines from ACI or visit ACI’s
website at www.concrete.org prior to
submitting contributions.
Papers reporting research must include
a statement indicating the significance of
the research.
The Institute reserves the right to return,
without review, contributions not meeting
the requirements of the Publication Policy.
All materials conforming to the Policy
requirements will be reviewed for editorial
quality and technical content, and every
effort will be made to put all acceptable
papers into the information channel.
However, potentially good papers may be
returned to authors when it is not possible
to publish them in a reasonable time.
Discussion
All technical material appearing in the
ACI Materials Journal may be discussed.
If the deadline indicated on the contents
page is observed, discussion can appear
in the designated issue. Discussion should
be complete and ready for publication,
including finished, reproducible illustra-
tions. Discussion must be confined to the
scope of the paper and meet the ACI Publi-
cation Policy.
Follow the style of the current issue.
Be brief—1800 words of double spaced,
typewritten copy, including illustrations
and tables, is maximum. Count illustrations
and tables as 300 words each and submit
them on individual sheets. As an approxi-
mation, 1 page of text is about 300 words.
Submit one original typescript on 8-1/2 x
11 plain white paper, use 1 in. margins,
and include two good quality copies of the
entire discussion. References should be
complete. Do not repeat references cited
in original paper; cite them by original
number. Closures responding to a single
discussion should not exceed 1800-word
equivalents in length, and to multiple
discussions, approximately one half of
the combined lengths of all discussions.
Closures are published together with
the discussions.
Discuss the paper, not some new or
outside work on the same subject. Use
references wherever possible instead of
repeating available information.
Discussion offered for publication should
offer some benefit to the general reader.
Discussion which does not meet this
requirement will be returned or referred to
the author for private reply.
Send manuscripts to:
http://mc.manuscriptcentral.com/aci
Send discussions to:
Journals.manuscripts@concrete.org
ON FRONT COVER: 118-M115, p. 395, Fig. 1—Gantry 3D-printing system and various
products (left); 118-M108, p. 329, Fig. 6—Simulation of suspension flow in six-blade-vane
rheometer. Only solids and impeller are shown; right image shows a later stage of flow. Note
migration of solids away from impeller in latter case (center); 118-M98, p. 205, Fig. 2—Wall
effect on formation of LL (right).
ON BACK COVER: 118-M82, p. 23, Fig. 4(a)—Buildability performance quantification via
printing circular hollow column.
167
Rheology and Setting Control of Concrete for Digital Construction, by
L. S. C. Ko, S. Moro, J. Bury, T. Vickers, B. Sachsenhauser, and S. Mönnig
177
Rheological Properties of Metakaolin-Based Geopolymers for Three-
Dimensional Printing of Structures, by M. D. M. Paiva, L. D. Fonseca
Rocha, L. I. Castrillon Fernandez, R. D. Toledo Filho, E. C. C. M. Silva,
R. Neumann, and O. A. Mendoza Reales
189
Rheological Characterization of Three-Dimensional-Printed Polymer
Concrete, by D. Heras Murcia, M. Abdellatef, M. Genedy, and M. M. Reda Taha
203
Computational Investigation of Concrete Pipe Flow: Critical Review,
by Tooran Tavangar, Masoud Hosseinpoor, Ammar Yahia, and Kamal H.
Khayat
217
Characterization of Tensile Behavior of Fresh Cementitious Materials,
by Y. Jacquet, V. Picandet, and A. Perrot
227
Concrete Shear Box: New Instrument to Assess Stiff to Flowing
Concrete Using Bingham Model, by Girish Shamanna, Ajay Nagaraj, and
Akanksha Achutha
241
Rheological Properties of Recycled Aggregate Concrete Incorporating
Supplementary Cementitious Materials, by Samer Al Martini, Ahmad
Khartabil, and Narayanan Neithalath
255
Reinforcement of Three-Dimensional-Printed Concrete Structures
by Weaving and Knitting of Continuous Fibers and Wires, by Gregor
Fischer and Ieva Paegle
263 Rheological Response of Magnetorheological Cementitious Inks Tuned
for Active Control in Digital Construction, by Aparna S. Deshmukh,
Reed T. Heintzkill, Rosalba A. Huerta, and Konstantin Sobolev
275
Use of Nanoclays and Methylcellulose to Tailor Rheology for Three-
Dimensional Concrete Printing, by AlaEddin Douba and Shiho Kawashima
291 Benefits of In-Transit Management Systems through Addition of Admix-
ture, by Jason Straka, Stephen P. Klaus, Junfeng Zhu, Pete A. Gentile, and
Nathan A. Tregger
301
Three-Dimensional (3D)-Printed Wood-Starch Composite as Support
Material for 3D Concrete Printing, by Viacheslav Markin, Christof Schröfl,
Paul Blankenstein, and Viktor Mechtcherine
311
How Admixtures Affect Yield Stresses of Cement, by Chandrasekhar
Bhojaraju, Malo Charrier, and Claudiane M. Ouellet-Plamondon
325
Standard Reference Materials for Rheological Measurements of
Cement-Based Materials, by N. S. Martys, W. L. George, S. G. Satterfield,
B. Toman, M. A. Peltz, S. Z. Jones, and C. F. Ferraris
331
Reinforcing Interlayers of Three-Dimensional-Printed Mortar Using
Metal Fiber Insertion, by Tomoya Nishiwaki, Yoshihiro Miyata, Shoko
Furue, Shiko Fukatsu, and Hideyuki Kajita
341
Effect of Particle Contact and High-Range Water-Reducing Admixture
Adsorption on Rheology of Cement Paste, by Jiang Zhu, Jiaping Liu,
Kamal H. Khayat, Xin Shu, and Zhen Li
353
Controlling Three-Dimensional-Printable Concrete with Vibration, by
Karthik Pattaje Sooryanarayana, Kathleen A. Hawkins, Peter Stynoski, and
David A. Lange
Contents continue on next page
@seismicisolation
@seismicisolation
4. ACI Materials Journal/November 2021 3
359
Estimating Rheological Properties of Superplasticized Cement Paste at High Temperature and Prolonged
Mixing Using Various Rheological Models and Oscillatory Rheology, by Samer Al Martini, Ahmad Khartabil, and
Moncef L. Nehdi
371
Linking Solids Content and Flow Properties of Mortars to their Three-Dimensional Printing Characteristics, by
Fabian B. Rodriguez, Jan Olek, Reza Moini, Pablo D. Zavattieri, and Jeffrey P. Youngblood
383
Impact of Chemical Admixtures on Time-Dependent Workability and Rheological Properties of Ultra-High-
Performance Concrete, by Flavia Mendonca and Jiong Hu
395
Case Study: Measuring Flow and Setting Time for Three-Dimensionally Printed Mortar, by Emily Xu, Karin
Vanessa Tejada, David Walker, Sasa Zivkovic, and Kenneth C. Hover
407
Testing and Modeling of Reinforcement Cage Penetration Capacity from Concrete Rheometer Results, by
Yannick Vanhove, Chafika Djelal, and Albert Magnin
421
Characterizing Sticky Concrete from Rheological Perspective, by Elizabeth G. Burns, Danila F. Ferraz, and Nathan
A. Tregger
431 Discussion
Tensile Properties of Basalt Fiber-Reinforced Polymer Reinforcing Bars for Reinforcement of Concrete. Paper by R. Kamp-
mann, S. Telikapalli, A. Ruiz Emparanza, A. Schmidt, and M. A. Dulebenets
LETTER FROM THE GUEST EDITORS
With the increasing use of flowable concrete, considerable
advances have been made in recent years to develop tools
to investigate the flow behavior of cement-based materials
and tailor specialty admixtures to control the rheology of
concrete. Several advances have been made to model the
flow behavior of concrete to solve complex problems related
to concrete materials behavior, prediction, and performance.
More recently, additive manufacturing, commonly known as
three-dimensional (3D) printing, has emerged as a revolu-
tionary technology that has contributed to advancements in
a wide range of industries, such as aerospace, automotive,
and healthcare. Exploiting the transformative impact of 3D
printing in construction requires the design and implemen-
tation of complex engineered materials that can meet the
versatile requirements of this technology.
The goal of this special issue was to ramp up the number
of impactful, high-quality manuscripts using these methods
in the ACI Materials Journal. While currently used primarily
by researchers, the broader ACI community should have a
particular interest in the importance of rheological proper-
ties on the performance specification of concrete that plays
a vital role in mixing, placement, consolidation, finishing,
segregation resistance, mechanical properties, and durability
and service life. The science of rheology and 3D-printing
technology can ultimately revolutionize the way concrete
design guides or manuals, standards, and codes are addressed
and implemented.
This special issue of the ACI Materials Journal contains
37 select manuscripts focused on rheology and 3D printing
of cement-based materials. The papers cover a wide range
of topics, including advances in characterization and testing
techniques and instrumentation to assess concrete rheology
and workability; effect of chemical admixtures (including
high-range water-reducing admixtures [HRWRAs],
retarders, and viscosity-modifying agents), mineral admix-
tures (such as nanoclays and supplementary cementitious
materials), binder compositions (for example, limestone
calcined clay and alkali-activated), fibers (metal and natural),
and system interactions (for example, HRWRA adsorption)
on the rheological properties of cement-based systems,
including 3D-printing concretes; and fundamental rheolog-
ical studies to facilitate concrete pumping and casting appli-
cations, including extrusion-based 3D concrete printing.
Kamal H. Khayat
Shiho Kawashima
@seismicisolation
@seismicisolation
9. 8 ACI Materials Journal/November 2021
viscosity and yield stress of fresh concrete.39,40
The pumping
pressure of fresh concrete using coarse aggregate with a
maximum particle size of 25 mm (1 in.) was twice that with
a maximum particle size of 10 mm (0.4 in.).41
The change
of aggregate particle size was accompanied by the change
of mixing water and HRWRA dosage. However, the differ-
ence in coarse aggregate particle size had no effect on the
thickness of the lubrication layer. Reinhardt et al.42
studied
the effect of the content and composition of aggregate on
the rheological behavior of SCC. The yield stress and plastic
viscosity of fresh concrete were described with the model
parameter of thickness of excess paste.42
The present study was designed to investigate the change
in rheological properties of the lubrication layer with the
change in aggregate content. Measurements of the tribom-
eter, sliding pipe rheometer, and mortar rheology were
employed to determine the yield stress and viscosity of the
lubrication layer. The performance indexes of fresh concrete,
including the rheological properties, thixotropy, slump flow,
and V-funnel flow time, were studied in parallel to compre-
hensively characterize the effect of aggregate content on the
performance of pumping concrete.
RESEARCH SIGNIFICANCE
The lubrication layer plays a governing role in predicting
the pumpability of fresh concrete. This study simultaneously
evaluated the effects of aggregate content on the rheological
properties of the lubrication layer and its formation with the
rheological properties of bulk concrete. The effects of aggre-
gate content on the amount of shearing in the bulk concrete
and the degree of SIPM were studied. Three different test
methods were used simultaneously to determine the rheo-
logical properties of the lubrication layer, and a critical
review of the pros and cons of each method was provided.
EXPERIMENTAL PROCEDURE
Materials
The P∙I 42.5 portland cement, fly ash, and silica fume,
which conformed to Chinese National Standard GB
175-2007 (equivalent to European CEM I 42.5), GB/T 1596,
and GB/T 27690, respectively, were used for the prepara-
tion of fresh concrete in this study. The density of these
raw materials was 3045 kg/m3
(190 lb/ft3
), 2363 kg/m3
(148 lb/ft3
), and 2242 kg/m3
(140 lb/ft3
), respectively, and
their chemical compositions are displayed in Table 1. A
commercially available polycarboxylate-based HRWRA
(PCE) with a solids content of 50% was used to adjust the
workability of concrete. Its chemical structure is shown in
Fig. 1. The characteristic properties of the PCE and its anionic
charge density can be seen in Table 2. Crushed limestone
with an apparent density of 2700 kg/m3
(169 lb/ft3
) and bulk
density of 1486 kg/m3
(93 lb/ft3
) was used as the coarse aggre-
gate. Its particle size was 5 to 10 mm (0.2 to 0.4 in.). The
maximum aggregate size is less than one-third of the radius
of the inner rotor and the distance between the inner rotor and
outer cylinder to ensure that the measured fresh concrete can
be regarded as a fluid in the rheometer. Natural river sand with
an apparent density of 2675 kg/m3
(167 lb/ft3
) and a fineness
modulus of 2.6 was used as the fine aggregate. Its bulk density
was 1425 kg/m3
(89 lb/ft3
). The particle size distribution of
coarse and fine aggregate is shown in Fig. 2.
Mixture design
Mixture proportions of concrete—Different mixture
proportions of concrete were designed as shown in Table 3.
The mixture proportion of reference sample C3 was based
on a commonly used mixture proportion of SCC. The
proportion of fly ash in the binder is appropriate, which
can further improve the workability and later strength of
concrete.43
Incorporating a proper proportion of silica fume
can improve the mechanical properties of concrete.44
To
explore the effect of aggregate content on the rheological
properties of the lubrication layer and fresh performance
of bulk concrete, two variables were adopted: sand-to-total
Table 1—Chemical compositions of cement, fly ash, and silica fume, w/%
Composition CaO SiO2 Al2O3 Fe2O3 MgO SO3 Na2Oeq
Cement 63.27 22.59 4.42 3.44 2.43 2.41 0.38
Fly ash 4.59 48.98 31.43 9.13 0.49 1.52 1.06
Silica fume 0.26 97.12 0.06 0.18 0.36 0.1 1.21
Note: Na2Oeq = Na2O + 0.658K2O.
Fig. 1—Chemical structure of PCE.
Table 2—Characteristic properties of PCE
Polymer
Mw,
g/mol
DPI,
Mw/
Mn
Mass fraction of polymer
in synthesized samples,
%
Anionic charge
density, µeq/g,
pH = 12
PCE 39,144 1.94 90.08 1191
@seismicisolation
@seismicisolation
10. 9
ACI Materials Journal/November 2021
aggregate ratio (sand ratio) and paste-aggregate ratio (PA
ratio). The sand ratio was used to characterize the effect
of the content ratio of coarse and fine aggregate when the
total aggregate amount was fixed. The paste-aggregate
ratio indicates the change in aggregate content in concrete.
In this study, the variation ranges of the sand ratio and PA
ratio were 45% to 57% and 44% to 88%, respectively. On
one hand, the values of sand ratio and PA ratio used in
some studies39,40
were within this range. On the other hand,
expanding the experimental range of sand ratio and PA ratio
can more comprehensively reflect the influence of aggregate
content on the rheological properties of lubrication layer and
evaluate different test methods. A water-binder ratio (w/b)
of 0.3 was used, which was commonly used in practical
engineering and research.44
The total amount of water in the
mixture proportions takes into account the water-absorbing
capacity of coarse aggregate and fine aggregate. It should
be noted that the concrete mixture design in this study was
narrow in scope with all mixture designed with the same w/b
and binder composition, and that the mixtures are not repre-
sentative of SCC or highly flowable concrete consistently.
Calculation of paste thickness on coated aggregate—
Assuming that the thickness of the excess paste on coated
aggregate is uniform and much smaller than the particle size
of aggregate, the total volume of paste in the concrete is45
V t M A M A V
paste paste F F C C void
(1)
where Vpaste is the unit volume of the paste (m3
); tpaste is the
average paste thickness on the coated aggregate (m); MF and
MC are the dosage of fine aggregate and coarse aggregate,
respectively (kg); AF and AC are the specific surface area of
fine aggregate and coarse aggregate respectively (m2
/kg);
and Vvoid is the volume of paste required to fill the pores
between the aggregate (m3
).
The specific surface area of fine aggregate and coarse
aggregate can be calculated by assuming that the aggregate
is a spherical particle, as shown in Eq. (2)
A
m
D
M
K
D
i
i
i
i
i
i
=
∑
= ∑
6
6
ρ
ρ
(2)
where A is the specific surface area of aggregate (m2
/kg); Di
is the intermediate particle size of each aggregate gradation
(m); M is the total mass of aggregate (kg); mi is the total
mass of each aggregate gradation (kg); and Ki is the mass
fraction of each aggregate gradations.
The void volume of the mixed aggregate Vvoid can be
calculated by Eq. (3) using the apparent density and bulk
density of the mixed aggregate
V
M M
void
A
A
A
A
(3)
where MA is the total mass of mixed aggregate (kg); ρA is
the bulk density of mixed aggregate (kg/m3
); and ρA′ is the
apparent density of mixed aggregate (kg/m3
).
The bulk density of mixed aggregate and the calculated
paste thickness on coated aggregate are shown in Table 3.
Fig. 2—Gradation curves of fine and coarse aggregate.
Table 3—Mixture proportions of concrete
Sample
Ingredient, kg/m3
PCE, %
Sand
ratio, %
PA ratio,
% w/b
Bulk density
of mixed
aggregate, kg/m3
Paste thickness
on coated
aggregate, μm
Cement Fly ash Silica fume Fine aggregate Coarse aggregate
C1 401 150 25 728 890 0.55 45 63 0.3 1760 20.1
C2 401 150 25 776 841 0.55 48 63 0.3 1772 20.8
C3 401 150 25 823 794 0.55 51 63 0.3 1796 23.1
C4 401 150 25 873 744 0.55 54 63 0.3 1768 18.4
C5 401 150 25 922 695 0.55 57 63 0.3 1758 16.3
C6 319 119 21 933 900 0.55 51 44 0.3 1796 0
C7 360 135 23 878 847 0.55 51 53 0.3 1796 7.2
C8 442 165 29 768 741 0.55 51 75 0.3 1796 41.7
C9 483 181 31 714 688 0.55 51 88 0.3 1796 62.9
Note: 1 kg/m3
= 0.0624 lb/ft3
.
@seismicisolation
@seismicisolation
11. 10 ACI Materials Journal/November 2021
Testing procedure
Mixing—First, the cementitious materials and fine aggre-
gate were added into a mixer, and the raw materials were
mixed for 2 minutes to ensure the homogeneity of mixture.
Then, half of the required water was added into the mixture,
and the material was mixed for 2 minutes. The remaining
water and PCE were subsequently added, and the mixture
was mixed for another 2 minutes. Simultaneously, the
internal surface of the mixer was manually scraped. The
coarse aggregate was finally added into the mixer and mixed
for 3 minutes. The total mixing time was sufficiently long to
ensure the homogeneity of the high-viscosity fresh concrete.
Due to the retarding effect of PCE, the workability of the
studied fresh concrete remained basically unchanged in
1 hour.
Rheological measurement of constitutive mortar—The
constitutive mortar, which was wet-screened from the fresh
concrete, was obtained and measured to determine the rheo-
logical properties of the lubrication layer. A coaxial rheom-
eter was used to complete the rheological test of the mortar.
A four-bladed vane with a height (h1) of 70 mm (2.8 in.) and
a radius (Ri) of 35 mm (1.4 in.) was used as the inner rotor.
The radius (Ro) of the outer cylinder was 82 mm (3.28 in.),
which was ribbed to prevent the materials from slipping. The
photographs of rheometer and inner rotor are shown in Fig. 3.
Figure 4 shows the scheme to test the rheological properties
of the mortar. First, the thixotropy was measured. The rota-
tional speed increased to 90 rpm within 2 minutes and subse-
quently linearly decreased to zero within 2 minutes. The area
of the thixotropic ring, which was surrounded by the torque
and rotational speed, was used to characterize the thixotropy
of the mortar.46
Then, the rheological test was performed.
The rotational speed increased to a maximum of 90 rpm and
remained for 45 seconds as a preshearing process, which can
minimize the effect of thixotropy. Then, it decreased step-
wise. Reliable testing results can only be obtained under
the condition of steady flow. To ensure the mortar reach a
stable flow state, the rotational speed in each platform of
the testing scheme was maintained for 15 seconds. The
Bingham model was used to represent the rheology model of
the measured mortar. A linear relationship was obtained by
fitting the acquired torque-rotational speed data. According
to the Reiner-Riwlin equation (Eq. (4)) and the solution of
Couette inverse problem, the rheological parameters of the
mortar were obtained by calculating the intercept and slope
of the linear relation
T
h
R R
h
R R
R
R
m
i o i o
o
i
m
=
−
+
−
4
1 1
4
1 1
1
2 2
1
2 2
π µ π
τ
Ω ln (4)
where τm is the yield stress of mortar (Pa); μm is the plastic
viscosity of mortar (Pa·s); T is the torque (N·m); and Ω is
the angular velocity (rad/s). Ro is replaced by plug radius Rp
when the mortar is partly sheared.
Rheological measurement with sliding pipe rheometer—
The sliding pipe rheometer shown in Fig. 5, which was
designed based on Kaplan et al. model, is the latest device
to evaluate the concrete pumpability.47
The rheological test
of the lubrication layer was performed with the sliding pipe
rheometer at the same time with the mortar measurement.
Before the test, the fresh concrete maintained stationary, and
the sliding pipe rheometer was preslid several times. Hence,
a lubrication layer was formed on the pipe wall. Multiple
counterweights were adopted, and each weight was slid three
times in the test process. The pressure-flow rate data can be
calculated from the curves of pressure and sliding velocity,
which were recorded with the pressure sensor and distance
sensor at each slide. According to the adopted model of
the sliding pipe rheometer, the relationship between pres-
sure and flow rate is shown in Eq. (5). The yield stress and
viscosity constant of the lubrication layer were obtained
from the intercept and slope of this linear relationship
Fig. 3—(a) Coaxial rheometer; and (b) inner rotor.
Fig. 4—Testing scheme to measure rheological parameters
of constitutive mortar.
@seismicisolation
@seismicisolation
12. 11
ACI Materials Journal/November 2021
P
l
d
lQ
d
s s
4 16
3
(5)
where P is the pressure (Pa); Q is the flow rate (m3
/s); τs
is the yield stress of the lubrication layer (Pa); ηs = μs/e is
the viscous constant of the lubrication layer (Pa·s/m); μs is
the plastic viscosity of the lubrication layer (Pa·s); and e is
the thickness of the lubrication layer (m). d = 0.126 m and
l = 0.5 m are the diameter and length of the sliding pipe
rheometer, respectively.
Rheological measurement of concrete—The rheological
properties of fresh concrete were determined with a coaxial
rheometer, which were required to calculate the rheolog-
ical parameters of the lubrication layer with the tribometer.
The distance between the outer cylinder and the inner rotor
was greater than three times the maximum particle size of
the coarse aggregate. The scheme to measure the rheolog-
ical parameters of fresh concrete was similar to the testing
scheme in Fig. 4. However, the maximum rotational speed
was 60 rpm. First, the thixotropy of fresh concrete was
measured. Then, the rheological test was performed, and
the stable rotational speed in each platform of the scheme
was maintained for 20 seconds. Yield stress τc and plastic
viscosity μc of the fresh concrete were calculated based on
Reiner-Riwlin equation (Eq. (4)).
Simultaneously, the concrete workability including the
V-funnel flow time and slump flow was measured.
Rheological measurement with the tribometer—The
tribometer is often used to study the rheological parame-
ters of the lubrication layer. The coaxial rheometer was also
used for this measurement. However, the inner vane rotor
was replaced by a steel cylinder with a radius of 35 mm
(1.4 in.), which is shown in Fig. 6. The surface of the inner
cylinder was smooth, and its bottom was conical, which was
designed for better insertion into concrete. The rheological
scheme in this measurement was identical to that of the
mortar test (Fig. 4). The thixotropy was not measured. More-
over, the preshearing process here was to fully form a lubri-
cation layer near the surface of the inner cylinder. During
actual pumping, the concrete was sheared at a high shear
rate to form a lubrication layer. Assuming that the thickness
of the lubrication layer was 2 mm (0.08 in.)11,14
and based on
Eq. (6), the maximum shear rate generated by the tribometer
in this test was approximated as 179 1/s, which satisfies the
shear rate in actual engineering. However, a change in the
assumed thickness of the lubrication layer directly affects
the calculated maximum shear rate. The increase of the
thickness of lubrication layer will lead to the decrease of the
shear rate
max
i
i l
R
R R
2
2 2
2
1 1
(6)
where Rl is the distance from the center of the inner cylinder
to the end of the lubrication layer in the radial direction (m);
and Ri is the radius of the inner cylinder (m).
Fig. 5—Sliding pipe rheometer.
Fig. 6—Inner rotor of tribometer.
@seismicisolation
@seismicisolation
13. 12 ACI Materials Journal/November 2021
Rheological measurements were performed at two
different filling heights to eliminate the bottom effect of
the cone part. The finally applied height (h2) was the differ-
ence between the two filling heights, which was 71 mm
(2.84 in.). The torque difference between the two tests was
the final torque value in the following calculation. Due to
the unknown thickness of the lubrication layer, the viscosity
constant was used to characterize its rheological properties
as shown in Eq. (7)
t t l i
R
(7)
where Ωl is the angular velocity (rad/s); τt is the yield stress
of the lubrication layer (Pa); and ηt is the viscous constant of
the lubrication layer (Pa·s/m). The shear stress can be trans-
formed from the torque at the inner cylinder by Eq. (8).
T
R h
i
2 2
2
(8)
The effect of concrete shearing on the angular velocity
should be eliminated to obtain the angular velocity contrib-
uted by the shear flow of the lubrication layer. The flowing
state of the bulk concrete must be first determined by calcu-
lating the plug radius with Eq. (9)48
R
T
h
p
c
2 2
(9)
where τc is the yield stress of concrete (Pa). When Rp is
smaller than radius Ri of the inner cylinder, the bulk concrete
is not sheared. When Rp is larger than radius Ro of the outer
cylinder, the bulk concrete is completely sheared. Other-
wise, the bulk concrete is partly sheared. The bulk concrete
is usually not sheared for concrete with high yield stress.
However, for HFC and SCC, which have small yield stress
values, or when the torque is sufficiently high, the shear
flow occurs in the bulk concrete. If the bulk concrete is fully
sheared, the angular velocity contributed by the shear flow
of concrete can be calculated with Eq. (10)
c
c i o
c
c
o
i
T
h R R
R
R
4
1 1
2
2 2
ln (10)
where μc is the viscosity of bulk concrete; and Ωc is the
angular velocity contributed by the shear flow of concrete
(rad/s). When the bulk concrete is partly sheared, Ro is
replaced by Rp in Eq. (10). If the bulk concrete is not sheared,
Ωc should be equal to zero.
For each imposed rotational speed, the angular velocity
was corrected by the equation of Ωl = Ω– Ωc, where Ω is the
angular velocity of the cylinder. Thus, the rheological prop-
erties of the lubrication layer were calculated with Eq. (7)
based on the corrected torque and angular velocity.
RESULTS AND DISCUSSION
Effect of sand ratio
Performance of fresh concrete—Figure 7 shows the change
in performance of fresh concrete, including its rheological
properties and workability, with the change in sand ratio. As
displayed in Fig. 7(a), the evolutions of yield stress τc and
plastic viscosity μc of fresh concrete show opposite trends.
τc first rapidly decreases and subsequently increases in the
range of sand ratio in this paper. In other words, τc reaches
a minimum at a certain sand ratio. The minimum yield
stress is almost half of the maximum. μc shows a tendency
to first increase and subsequently decrease. However, in the
moderate range of sand ratio, the change in μc is small. Once
the sand ratio is too high or too low, μc significantly changes.
It has been accepted that the yield stress is the minimum
shear stress required for the flow of concrete, which is
generated by the cohesive network structure of solid parti-
cles and the adhesion and friction among the particles.49
The
plastic viscosity is formed by the interaction force of parti-
cles and viscous force among particles, which reflects the
rate of deformation of fresh concrete under shear stress.50
According to the theory of excess paste, the cementitious
paste in fresh concrete consists of two parts: one of which is
used to fill the voids among the aggregate particles, and the
other part wraps the surface of aggregate particles to provide
fluidity. The change in sand ratio indicates the change in
concentration of coarse and fine particles and the packing
state of particles. For concrete with low sand ratio, the voids
among coarse aggregates are originally filled with fine aggre-
gates and partly paste, which results in a decrease in thick-
ness of the paste layer that wraps the aggregate particles.
When the sand ratio is high, the increase in fine aggregate
concentration increases the surface area of the aggregate,
Fig. 7—Evolution of: (a) rheological properties; and (b) workability of concrete with different sand ratios. (Note: 1 Pa =
0.000145 psi; 1 mm = 0.039 in.)
@seismicisolation
@seismicisolation
14. 13
ACI Materials Journal/November 2021
which also decreases the thickness of the paste layer that
wraps the aggregate particles. The reduction of the paste
layer enhances the friction and interlocking effect among
the aggregate particles, which increases the yield stress of
concrete. The close packing of aggregates helps to reduce the
yield stress of concrete. With the increase in sand ratio, the
content of fine particles and powder in aggregate increases,
which improves the cohesiveness of the paste. Therefore, the
plastic viscosity of concrete gradually increases. However,
the settlement of coarse aggregate occurs when the sand
ratio is high, which may decrease the plastic viscosity and
increase the yield stress.
Although no fundamental physical parameters are
provided, the workability indexes of fresh concrete, such
as the slump flow and V-funnel flow time, are frequently
used on site to estimate the concrete pumpability.51
It has
been accepted that the slump flow and V-funnel flow time
of concrete have a good correlation with its yield stress
and plastic viscosity, respectively. The change in work-
ability of fresh concrete is shown in Fig. 7(b). The slump
flow and V-funnel flow time of concrete first increase and
subsequently decrease. This result verifies the relationship
between rheological properties and workability of concrete.
Meanwhile, a moderate sand ratio can improve the fluidity
of concrete but slightly increase its cohesiveness.
Rheological properties of lubrication layer and pumpa-
bility of concrete—The change in rheological properties of
the lubrication layer is determined by different measurement
systems with the change in sand ratio as shown in Fig. 8.
Although there is a difference in shear rate and moving state
of concrete during the measurement, the sliding pipe rheom-
eter and tribometer have similar testing principles. The
SIPM and wall effect result in the formation of the lubrica-
tion layer. For the measurement with sliding pipe rheometer,
yield stress τs of the lubrication layer rapidly decreases with
the increase in sand ratio and subsequently increases, while
viscous constant ηs of the lubrication layer first increases
and subsequently decreases. The change in τs is obvious,
but ηs only slightly changes in the case of three moderate
sand ratios. The tribometer obtains a similar varying trend of
the rheological properties of the lubrication layer, while no
obvious inflection point is observed for yield stress τt. More-
over, the absolute values of the parameters obtained with the
two methods are significantly different. Some studies have
shown that the yield stress of the lubrication layer was very
small and even zero, which was far lower than the yield
stress of bulk concrete.34
However, some research pointed
out that the yield stress of lubrication layer may be very
large, and even exceed that of concrete.47
This may depend
on the measurement method, which will be discussed later.
Some studies have demonstrated that particles migrate
from the highly sheared zones due to particle collisions.33
The
competition between gradients in viscosity of the suspension
and gradients in particle collision frequency causes the ulti-
mate balance of SIPM. The migration coefficient of parti-
cles is proportional to the square of the particle size, shear
rate, and square of the solid volume fraction.13,52
Therefore,
the coarse aggregate particles and some sand grains migrate
under the effect of SIPM until the viscosity of the destination
zone, to which the particles are migrating, is sufficiently
high, and the local particle volume fraction conforms to the
random loose packing. With the initial increase in sand ratio,
the volume of coarse aggregate decreases, and the volume
of fine aggregate increases, which changes the bulk density
of particles and increases the volume of the paste layer on
the surface of aggregate particles. The migration of coarse
Fig. 8—Evolution of rheological properties of lubrication
layer with different sand ratios determined by: (a) sliding
pipe rheometer; (b) tribometer; and (c) mortar rheology.
(Note: 1 Pa = 0.000145 psi; 1 Pa·s/m = 3.68 × 10–6
psi·s/in.)
@seismicisolation
@seismicisolation
15. 14 ACI Materials Journal/November 2021
particles makes more paste or mortar migrate to the pipe wall.
The yield stress of the lubrication layer decreases. However,
the viscous constant of the lubrication layer is a comprehen-
sive index, which is related to the viscosity and thickness of
the lubrication layer. The increase in viscosity or decrease
in thickness can improve the viscous constant. With further
increase in sand ratio, the migration of coarse aggregate
particles intensifies. However, the paste layer on the surface
of aggregate particles is reduced. The bulk concrete can be
sharply sheared, and the formation of a lubrication layer is
affected, which will be discussed in the following. There-
fore, the yield stress and viscous constant of the lubrication
layer have different inflection points from concrete.
As mentioned, the concrete pumpability depends on the
rheological properties of the lubrication layer. For HFC and
bulk concrete, which is sheared during pumping, Kaplan
et al. model based on Bingham model is often used to predict
the concrete pumpability,29
as shown in Eq. (11)
P
L
D
Q
D
D D
D
gh
c
LL
c
c
c
LL
LL LL
4 900 8 6
1
8
2
(11)
where D is the diameter of the pipe (m); L is the pipe length
(m); τLL is the yield stress of the lubrication layer (Pa); ηLL
is the viscous constant of the lubrication layer (Pa·s/m); α is
the filling coefficient of concrete; ρ is the density of concrete
(kg/m3
); and h is the pumping height (m).
According to the Kaplan et al. model and the measured
rheological properties of the lubrication layer with the
sliding pipe rheometer, when the sand ratio increases, the
required pumping pressure of concrete decreases at the same
flow rate. There is an optimal sand ratio that makes the best
pumpability of concrete. Although the viscous constant
of the lubrication layer is reduced at high sand ratio, its
yield stress improves, and the required pumping pressure
of concrete increases when the flow rate is not sufficiently
high. However, according to the test results with the tribom-
eter, the optimal sand ratio that yields the best pumpability
of concrete has not been observed, possibly due to different
shear states of concrete in the two test systems and the
limited range of sand ratio in this study.
The ratio of the viscous constant of the lubrication layer to
the viscosity of the bulk concrete (ηt/μc) is often used to char-
acterize the amount of shearing in the bulk concrete and the
degree of SIPM.10
Higher ηt/μc indicates an increase in shear
rate and sheared zone in the bulk concrete. The dilatancy
caused by the sheared bulk concrete pushes the coarse parti-
cles back toward the pipe wall or inner cylinder and prevents
their maximum packing.18
Thus, the formation of the lubri-
cation layer becomes difficult and less important. Figure 9
shows the change in ηt/μc of different concrete samples. The
ηt/μc of concrete with a moderate sand ratio is low, which
indicates that the formation of its lubrication layer during
pumping is easy. In the case with a high or low sand ratio,
ηt/μc rapidly increases. More bulk concrete is sheared, which
prevents the migration of coarse particles and formation of
the lubrication layer. The flowing state of fresh concrete in
the pump pipe is different under different sand ratios.
Compared with the test results of the sliding pipe rheom-
eter and tribometer, the mortar rheology test yields a different
developing trend of the rheological properties of the lubrica-
tion layer including yield stress τm and plastic viscosity μm.
τm slightly decreases at low sand ratio. When the sand ratio
continues to increase, τm gradually increases. The absolute
valueofτm isdifferentfromothertestresults.μm alsoincreases
with the increase in sand ratio. As mentioned, the formation
process, composition, and thickness of the lubrication layer
vary under different sand ratios. Although the lubrication
layer can be considered a thin layer of mortar, its maximum
size of particles and content of fine aggregate and paste are
affected by the bulk concrete and shearing process. However,
constitutive mortar is obtained by the wet-screening method,
which limits the maximum particle size and paste content.
The dynamic change of the lubrication layer is not fully
characterized. The thickness of the lubrication layer cannot
be considered with the use of plastic viscosity. Therefore, the
variation trend of the rheological properties of the lubrica-
tion layer obtained by the mortar rheology is different from
that with other measurement systems.
Fresh concrete is a type of thixotropic fluid. It is important
to determine the thixotropy of concrete in pumping applica-
tions. The use of fresh concrete with high thixotropy means
that the pumping pressure required in the initial stage of
concrete pumping is higher. The pumping pressure required
for concrete at the same flow rate gradually decreases as the
pumping process progresses. Therefore, understanding the
thixotropy of concrete is beneficial to control the pumping
process of concrete. Due to the effect of coarse and fine aggre-
gate, the thixotropy of fresh concrete is different from that of
paste. The thixotropy of fresh concrete and mortar is shown
in Fig. 10. The thixotropy of concrete and mortar gradually
increases with the increase in sand ratio, and the thixotropy
of concrete is higher than that of mortar. However, the thixo-
tropy of concrete rapidly decreases when the sand ratio is too
high, which is almost identical to that of mortar. Many factors
Fig. 9—Evolution of ηt/μc with different sand ratios. (Note:
1 m = 39.37 in.)
@seismicisolation
@seismicisolation
16. 15
ACI Materials Journal/November 2021
affect the thixotropy of concrete, including the total amount of
powder in the mixture, w/b, fineness of powder, and dosage of
the HRWRA.53
With the increase in sand ratio, for concrete,
due to the decrease in coarse aggregate content and lubrica-
tion effect of the increased paste layer on the surface of coarse
aggregate particles, the interlock effect of coarse aggregate is
reduced. However, the increase in powder content enhances
flocculation, and the increase in fine aggregate enhances
the interlock, which ultimately increases the thixotropy of
concrete. For mortar, although the proportion of cementitious
materials is reduced, the increase in flocculation caused by the
fine powder in aggregate and enhanced interlock of aggregate
particles also enhance the thixotropy. When the sand ratio is
too high, the content of coarse aggregate sharply decreases.
Thus, concrete has a similar thixotropy to mortar. Moreover,
the proportion of cementitious materials in the mortar is
further reduced, which weakens its thixotropy.
Effect of PA ratio
Performance of fresh concrete—The effect of the PA
ratio on the performance of fresh concrete is determined,
as shown in Fig. 11. The yield stress of concrete τc rapidly
decreases with the initial increase in PA ratio. With further
increase in PA ratio, τc decreases slightly. The minimum of
τc is almost one-fifth of the maximum. The plastic viscosity
of concrete μc has a similar evolution to τc. The increase
in PA ratio indicates the decrease in aggregate volume and
increase in paste volume, which directly increases the thick-
ness of the paste layer on the surface of aggregate particles,
as shown in Table 3. The lubrication effect of the paste layer
is enhanced, and the friction among aggregate particles is
reduced. Therefore, τc and μc of concrete decrease. However,
when the PA ratio is sufficiently high, concrete segregation
occurs. Although the paste thickness on coated aggregate
increases significantly in uniform concrete, the settlement of
the aggregate causes the measured yield stress of concrete
to be higher.
As shown in Fig. 11(b), the slump flow and V-funnel flow
time of fresh concrete show a good correlation with its yield
stress and plastic viscosity, respectively. The slump flow
of concrete rapidly increases with the increase in PA ratio
but slightly increases when its PA ratio exceeds 70%. The
V-funnel flow time of concrete reduces with the increase
in PA ratio. At the minimum PA ratio, the slump flow of
concrete is only 280 mm (11 in.), and there are obvious
coarse aggregates exposed on the surface of concrete. The
paste thickness on coated aggregate is 0. The concrete
cannot pass through the V-funnel. At the maximum PA ratio,
although the slump flow is large and the V-funnel flow time
is short, the concrete segregation is obvious. Therefore, the
PA ratio should be maintained in a certain range to optimize
the workability of concrete.
Rheological properties of lubrication layer and concrete
pumpability—The effect of the total amount of aggregate on
the performance of the lubrication layer is important because
it determines the amount of paste that can migrate to the lubri-
cation layer. The development of rheological properties of the
lubrication layer, which is measured by different methods,
with the change in PA ratio of concrete is shown in Fig. 12.
Yield stress τs and viscous constant ηs of the lubrication layer
measured by the sliding pipe rheometer generally decrease
with the increase in PA ratio. For the results determined with
the tribometer, τt and ηt gradually decreases with the increase
in PA ratio. However, the decrease of τt is more obvious when
PA ratio is higher. The absolute values of the test results with
two methods are different. The increase in PA ratio indicates
that more aggregate particles migrate in the radial direction
Fig. 10—Evolution of thixotropy of fresh concrete and
mortar with different sand ratios. (Note: 1 Nmrad/s =
8.85 lbf·in.·rad/s.)
Fig. 11—Evolution of: (a) rheological properties; and (b) workability of concrete with different PA ratios. (Note: 1 Pa =
0.000145 psi; 1 mm = 0.039 in.) @seismicisolation
@seismicisolation
17. 16 ACI Materials Journal/November 2021
away from the pipe wall or inner cylinder. The paste content
in the lubrication layer and its paste thickness increase, which
result in the decreased yield stress and viscous constant. Mean-
while, as shown in Fig. 13, ηt/μc gradually increases with the
increase in PA ratio, which indicates that more bulk concrete
is sheared. The dilatancy caused by the sheared bulk concrete
makes it difficult to form a lubrication layer. Therefore, the
rheological properties of the lubrication layer vary with the
change in PA ratio. The pumpability of concrete depends
on the combination of yield stress and viscous constant of
the lubrication layer. The concrete pumpability generally
improves with the increase in PA ratio.
The change in rheological properties of the constitutive
mortar is similar to that of the lubrication layer, which was
measured with the other two methods. Yield stress τm and
plastic viscosity μm continue to decrease with the increase in PA
ratio. The reduction of fine aggregate in the mortar due to the
increase in PA ratio directly causes the decrease in τm and μm.
Figure 14 shows the change in thixotropy of fresh concrete
and mortar, which decreases with the increase in PA ratio. At
the minimum PA ratio, the thixotropy of concrete is approx-
imately 3.5 times that of mortar and approximately 5 times
that of mortar at the maximum PA ratio. When the PA ratio
increases, the thixotropy of concrete and mortar rapidly
decreases. With further increase in PA ratio, the magnitude of
the reduction in thixotropy decreases. The lubrication effect of
paste on the surface of aggregate particles is enhanced when
the amount of coarse and fine aggregates reduces, which is
more obvious than the increased flocculation effect among
binder particles. This change results in the decrease in thixot-
ropy of concrete and mortar. The interlocking effect of aggre-
gate particles is weakened when the amount of aggregate is
too low. The opposite effects of lubrication and flocculation
continue to reduce the thixotropy but to a lesser extent.
In general, according to the formation principle of the
lubrication layer, the use of sliding pipe rheometer, tribom-
eter, and mortar rheometer are all effective methods to
determine the rheological properties of the lubrication layer.
Theoretically, the calculation results should be consistent
Fig. 12—Evolution of rheological properties of lubrication
layer with different PA ratios determined by: (a) sliding pipe
rheometer; (b) tribometer; and (c) mortar rheology. (Note:
1 Pa = 0.000145 psi; 1 Pa·s/m = 3.68 × 10–6
psi·s/in.)
Fig. 13—Evolution of ηt/μc with different PA ratios. (Note:
1 m = 39.37 in.)
Fig. 14—Evolution of thixotropy of fresh concrete and
mortar with different PA ratios. (Note: 1 m = 39.37 in.)
@seismicisolation
@seismicisolation
18. 17
ACI Materials Journal/November 2021
when measuring the same lubrication layer. However, as
shown previously, the absolute values of rheological param-
eters including yield stress and plastic viscosity are different.
In fact, although the principle of the various methods is
correct, the distinguishing formation processes of the lubri-
cation layer mean that the test objects are not identical. For
the sliding pipe rheometer, it can be used to quickly eval-
uate the pumpability of fresh concrete in the engineering
site. However, the effective length of the pipe is very short,
and the sliding speed of the pipe is slow especially when
the viscosity of concrete is high. Thus, the concrete shear
in a sliding pipe rheometer is different from that in a real
pump pipe. Moreover, it is found that the rubber ring in the
middle of the sliding pipe rheometer has a great effect on the
experimental results. Friction occurs when the rubber ring
is in close contact with the pipe, which prevents the pipe
from sliding. Conversely, when the contact between the two
is relatively loose, the mortar can leak from the gap because
of its small particle size, resulting in additional friction
during the slide of the pipe. The friction caused by the rubber
ring or the leaking mortar during the sliding pipe rheom-
eter measuring results in a significantly overestimated yield
stress of lubrication layer. Compared with the sliding pipe
rheometer, the formation of the lubrication layer during the
test with the tribometer is more similar to that in the actual
pump pipe. One of the main causes of error is the segrega-
tion. The torque at the bottom of the inner cylinder under
different filling heights may thus be various, resulting in the
deviation of the corrected torque. The maximum particle
size in the paste is limited when the rheological properties
of the lubrication layer are measured by mortar rheometer.
The migration process of particles and the thickness change
of lubrication layer will not be reproduced. The difference in
the formation of the lubrication layer results in the system-
atic deviation between different methods. It may be more
accurate to use a tribometer to characterize the rheological
properties of the lubrication layer.
CONCLUSIONS
1. The aggregate content has an important effect on the
rheological properties of the lubrication layer. With the
increase in sand ratio, the viscous constant of the lubrication
layer first increases and subsequently decreases. However,
the yield stress of the lubrication layer is rapidly reduced
and subsequently increases or continues to decrease, which
depends on the measurement system. The sliding pipe rheom-
eter test proves that there is an optimal sand ratio to yield
the best rheological properties of the lubrication layer and
concrete pumpability. The yield stress and viscous constant of
the lubrication layer are generally reduced with the increase
in paste-aggregate (PA) ratio. Therefore, the pumpability of
concrete is improved with the increase in PA ratio.
2. The changes of sand ratio and PAratio affect the particle
packing and paste thickness on coated aggregate, which also
affects the SIPM. There is an optimal sand ratio that mini-
mizes the shear of the bulk concrete. The shearing degree of
bulk concrete increases with the increase in PA ratio.
3. The yield stress of concrete is minimal at a moderate sand
ratio, while its plastic viscosity, V-funnel flow time, and slump
flow are maximal. The yield stress and plastic viscosity of
concrete decrease with the increase in PA ratio, and the work-
ability of concrete is improved. The thixotropy of concrete
first increases and subsequently decreases with the increase in
sand ratio, and it decreases when the PA ratio increases.
4. The formation process and composition of the lubrica-
tion layer in different test systems are varied, so the evolution
of the rheological properties of the lubrication layer varies.
The rheological test of mortar cannot fully characterize the
packing and migration process of particles when the sand
ratio changes, which leads to the great difference in rheo-
logical properties of the lubrication layer compared to the
results obtained by the sliding pipe rheometer or tribometer.
AUTHOR BIOS
Yu Liu is a Doctoral Candidate in the Department of Civil Engineering,
Tsinghua University, Beijing, China. He received his BE from the Univer-
sity of Science and Technology Beijing, Beijing, China. His research inter-
ests include rheology of cement-based materials, concrete pumping, and
self-consolidating concrete.
Rui Jing is an Assistant Professor in the Department of Civil Engineering,
Tsinghua University. He received his PhD from Tianjin University, Tianjin,
China. His research interests include self-consolidating concrete and
high-performance concrete.
Fengze Cao is a Doctoral Candidate in the Department of Civil Engi-
neering, Tsinghua University, where he also received his BE. His research
interests include expansive concrete and concrete pumping.
Peiyu Yan is a Professor in the Department of Civil Engineering, Tsinghua
University. He received his PhD from Wuhan University of Technology,
Wuhan, China. His research interests include cement-based materials and
concrete pumping.
ACKNOWLEDGMENTS
The authors would like to acknowledge National Key RD Program of
China (2017YFB0310101) and National Natural Science Foundation of
China (No. 51878381).
REFERENCES
1. Kwon, S. H.; Jang, K. P.; Kim, J. H.; and Shah, S. P., “State of the
Art on Prediction of Concrete Pumping,” International Journal of Concrete
Structures and Materials, V. 10, No. 3, 2016, pp. S75-S85. doi: 10.1007/
s40069-016-0150-y
2. Secrieru, E.; Cotardo, D.; Mechtcherine, V.; Lohaus, L.; Schroefl, C.;
and Begemann, C., “Changes in Concrete Properties during Pumping and
Formation of Lubricating Material under Pressure,” Cement and Concrete
Research, V. 108, 2018, pp. 129-139. doi: 10.1016/j.cemconres.2018.03.018
3. Feys, D.; Khayat, K. H.; and Khatib, R., “How do Concrete Rheology,
Tribology, Flow Rate and Pipe Radius Influence Pumping Pressure?”
Cement and Concrete Composites, V. 66, 2016, pp. 38-46. doi: 10.1016/j.
cemconcomp.2015.11.002
4. Kaplan, D.; de Larrard, F.; and Sedran, T., “Avoidance of Blockages
in Concrete Pumping Process,” ACI Materials Journal, V. 102, No. 3,
May-June 2005, pp. 183-191.
5. Vosahlik, J.; Riding, K. A.; Feys, D.; Lindquist, W.; Keller, L.; Van
Zetten, S.; and Schulz, B., “Concrete Pumping and Its Effect on the Air Void
System,” Materials and Structures, V. 51, No. 4, 2018, p. 94. doi: 10.1617/
s11527-018-1204-1
6. Kim, J. H.; Han, S. H.; and Choi, B. I., “Influence of Pumping Pressure
on the Viscosity Curve and Rheological Stability of Mortar Incorporating
Polycarboxylate,” Cement and Concrete Composites, V. 105, 2020, p. 7.
doi: 10.1016/j.cemconcomp.2019.103419
7. Rio, O.; Rodriguez, A.; Nabulsi, S.; and Alvarez, M., “Pumping
Quality Control Method Based on Online Concrete Pumpability Assess-
ment,” ACI Materials Journal, V. 108, No. 4, July-Aug. 2011, pp. 423-431.
8. Feys, D.; De Schutter, G.; Khayat, K. H.; and Verhoeven, R., “Changes
In Rheology Of Self-Consolidating Concrete Induced By Pumping,” Mate-
rials and Structures, V. 49, No. 11, 2016, pp. 4657-4677. doi: 10.1617/
s11527-016-0815-7
9. Choi, M. S.; Kim, Y. S.; Kim, J. H.; Kim, J.-S.; and Kwon, S. H.,
“Effects of an Externally Imposed Electromagnetic Field on the Formation
@seismicisolation
@seismicisolation
19. 18 ACI Materials Journal/November 2021
of a Lubrication Layer in Concrete Pumping,” Construction and Building
Materials, V. 61, 2014, pp. 18-23. doi: 10.1016/j.conbuildmat.2014.02.071
10. Feys, D., “How Much is Bulk Concrete Sheared during Pumping?”
Construction and Building Materials, V. 223, 2019, pp. 341-351. doi:
10.1016/j.conbuildmat.2019.06.224
11. Choi, M.; Roussel, N.; Kim, Y.; and Kim, J., “Lubrication Layer
Properties during Concrete Pumping,” Cement and Concrete Research,
V. 45, 2013, pp. 69-78. doi: 10.1016/j.cemconres.2012.11.001
12. Phillips, R. J.; Armstrong, R. C.; Brown, R. A.; Graham, A. L.; and
Abbott, J. R., “A Constitutive Equation for Concentrated Suspensions that
Accounts for Shear-Induced Particle Migration,” Physics of Fluids. A, Fluid
Dynamics, V. 4, No. 1, 1992, pp. 30-40. doi: 10.1063/1.858498
13. Leighton, D., and Acrivos, A., “A Constitutive Equation for Concen-
trated Suspensions that Accounts for Shear-Induced Particle Migration,”
Journal of Fluid Mechanics, V. 177, 1987, pp. 109-131. doi: 10.1017/
S0022112087000880
14. Choi, M. S.; Kim, Y. J.; and Kwon, S. H., “Prediction on Pipe
Flow of Pumped Concrete Based on Shear-Induced Particle Migration,”
Cement and Concrete Research, V. 52, 2013, pp. 216-224. doi: 10.1016/j.
cemconres.2013.07.004
15. Ngo, T. T.; Kadri, E. H.; Bennacer, R.; and Cussigh, F., “Use of Tribom-
eter to Estimate Interface Friction and Concrete Boundary Layer Composition
during the Fluid Concrete Pumping,” Construction and Building Materials,
V. 24, No. 7, 2010, pp. 1253-1261. doi: 10.1016/j.conbuildmat.2009.12.010
16. Ngo, T. T.; Kadri, E. H.; Cussigh, F.; Bennacer, R.; and Duval,
R., “Practical Tribometer to Estimate Pumpability of Fresh Concrete,”
Journal of Asian Architecture and Building Engineering, V. 9, No. 1, 2010,
pp. 229-236. doi: 10.3130/jaabe.9.229
17. Kwon, S. H.; Park, C. K.; Jeong, J. H.; Jo, S. D.; and Lee, S. H.,
“Prediction of Concrete Pumping: Part I—Development of New Tribometer
for Analysis of Lubricating Layer,” ACI Materials Journal, V. 110, No. 6,
Nov.-Dec. 2013, pp. 647-655.
18. Feys, D.; Khayat, K. H.; Perez-Schell, A.; and Khatib, R.,
“Development of a Tribometer to Characterize Lubrication Layer Proper-
ties of Self-Consolidating Concrete,” Cement and Concrete Composites,
V. 54, 2014, pp. 40-52. doi: 10.1016/j.cemconcomp.2014.05.008
19. Feys, D.; Khayat, K. H.; Perez-Schell,A.; and Khatib, R., “Prediction
of Pumping Pressure by Means of New Tribometer for Highly-Workable
Concrete,” Cement and Concrete Composites, V. 57, 2015, pp. 102-115.
doi: 10.1016/j.cemconcomp.2014.12.007
20. Nerella, V. N., and Mechtcherine, V., “Virtual Sliding Pipe Rheometer
for Estimating Pumpability of Concrete,” Construction and Building Mate-
rials, V. 170, 2018, pp. 366-377. doi: 10.1016/j.conbuildmat.2018.03.003
21. Secrieru, E.; Butler, M.; and Mechtcherine, V., “Testing Pumpability
of Concrete - From Laboratory Into Practice,” Bautechnik, V. 91, No. 11,
2014, pp. 797-811. doi: 10.1002/bate.201400072
22. Choi, M. S.; Park, S. B.; and Kang, S.-T., “Effect of the MineralAdmix-
tures on Pipe Flow of Pumped Concrete,” Journal of Advanced Concrete
Technology, V. 13, No. 11, 2015, pp. 489-499. doi: 10.3151/jact.13.489
23. Morinaga, S., “Pumpability of Concrete and Pumping Pressure in
Pipelines,” Proceedings of RILEM Seminar, Leeds, UK, 1973, pp 1-39.
24. Jacobsen, S.; Haugan, L.; Hammer, T. A.; and Kalogiannidis,
E., “Flow Conditions of Fresh Mortar and Concrete in Different Pipes,”
Cement and Concrete Research, V. 39, No. 11, 2009, pp. 997-1006. doi:
10.1016/j.cemconres.2009.07.005
25. De Schutter, G., and Feys, D., “Pumping of Fresh Concrete: Insights
and Challenges,” RILEM Technical Letters, V. 1, 2016, pp. 76-80. doi:
10.21809/rilemtechlett.2016.15
26. Le, H. D.; Kadri, E. H.; Aggoun, S.; Vierendeels, J.; Troch, P.; and De
Schutter, G., “Effect of Lubrication Layer on Velocity Profile of Concrete in
a Pumping Pipe,” Materials and Structures, V. 48, No. 12, 2015, pp. 3991-
4003. doi: 10.1617/s11527-014-0458-5
27. Secrieru, E.; Mohamed, W.; Fataei, S.; and Mechtcherine, V.,
“Assessment and Prediction of Concrete Flow and Pumping Pressure in
Pipeline,” Cement and Concrete Composites, V. 107, 2020, p. 103495. doi:
10.1016/j.cemconcomp.2019.103495
28. Chen, J.; Xie, H.; Guo, J.; Chen, B.; and Liu, F., “Preliminarily
Experimental Research on Local Pressure Loss of Fresh Concrete during
Pumping,” Measurement, V. 147, 2019, p. 106897. doi: 10.1016/j.
measurement.2019.106897
29. Kaplan, D.; de Larrard, F.; and Sedran, T., “Design of Concrete Pumping
Circuit,” ACI Materials Journal, V. 102, No. 2, Mar.-Apr. 2005, pp. 110-117.
30. Kwon, S. H.; Park, C. K.; Jeong, J. H.; Jo, S. D.; and Lee, S. H., “Predic-
tion of Concrete Pumping: Part II-Analytical Prediction and Experimental Veri-
fication,” ACI Materials Journal, V. 110, No. 6, Nov.-Dec. 2013, pp. 657-667.
31. Jeong, J. H.; Jang, K. P.; Park, C. K.; Lee, S. H.; and Kwon, S. H.,
“Effect of Admixtures on Pumpability for High-Strength Concrete,” ACI
Materials Journal, V. 113, No. 3, May-June 2016, pp. 323-333. doi:
10.14359/51688644
32. Jang, K. P., and Choi, M. S., “How Affect the Pipe Length of
Pumping Circuit on Concrete Pumping,” Construction and Building Mate-
rials, V. 208, 2019, pp. 758-766. doi: 10.1016/j.conbuildmat.2019.03.023
33. Spangenberg, J.; Roussel, N.; Hattel, J. H.; Stang, H.; Skocek, J.; and
Geiker, M. R., “Flow Induced Particle Migration in Fresh Concrete: Theo-
retical Frame, Numerical Simulations and Experimental Results on Model
Fluids,” Cement and Concrete Research, V. 42, No. 4, 2012, pp. 633-641.
doi: 10.1016/j.cemconres.2012.01.007
34. Kim, J. S.; Kwon, S. H.; Jang, K. P.; and Choi, M. S., “Concrete
Pumping Prediction Considering Different Measurement of the Rheo-
logical Properties,” Construction and Building Materials, V. 171, 2018,
pp. 493-503. doi: 10.1016/j.conbuildmat.2018.03.194
35. Feys, D.; De Schutter, G.; and Verhoeven, R., “Parameters Influ-
encing Pressure during Pumping of Self-Compacting Concrete,” Mate-
rials and Structures, V. 46, No. 4, 2013, pp. 533-555. doi: 10.1617/
s11527-012-9912-4
36. Jiao, D.; Shi, C.; Yuan, Q.; An, X.; Liu, Y.; and Li, H., “Effect of
Constituents on Rheological Properties of Fresh Concrete—A Review,”
Cement and Concrete Composites, V. 83, 2017, pp. 146-159. doi: 10.1016/j.
cemconcomp.2017.07.016
37. Aissoun, B. M.; Hwang, S.-D.; and Khayat, K. H., “Influence of
Aggregate Characteristics on Workability of Superworkable Concrete,”
Materials and Structures, V. 49, No. 1-2, 2016, pp. 597-609. doi: 10.1617/
s11527-015-0522-9
38. Tattersall, G. H., Workability and Quality Control of Concrete,
EFN Spon, London, UK, 1990.
39. Hu, J., and Wang, K., “Effect of Coarse Aggregate Characteristics on
Concrete Rheology,” Construction and Building Materials, V. 25, No. 3,
2011, pp. 1196-1204. doi: 10.1016/j.conbuildmat.2010.09.035
40. Hu, J., “A Study of Effects of Aggregate on Concrete Rheology,”
PhD dissertation, Iowa State University, Ames, IA, 2005.
41. Choi, M. S.; Kim, Y. J.; Jang, K. P.; and Kwon, S. H., “Effect of
the Coarse Aggregate Size on Pipe Flow of Pumped Concrete,” Construc-
tion and Building Materials, V. 66, 2014, pp. 723-730. doi: 10.1016/j.
conbuildmat.2014.06.027
42. Reinhardt, H. W., and Wuestholz, T., “About the Influence of the
Content and Composition of the Aggregates on the Rheological Behaviour
of Self-Compacting Concrete,” Materials and Structures, V. 39, No. 7,
2006, pp. 683-693. doi: 10.1617/s11527-006-9102-3
43. Uysal, M., and Sumer, M., “Performance of Self-Compacting
Concrete Containing Different Mineral Admixtures,” Construction and
Building Materials, V. 25, No. 11, 2011, pp. 4112-4120. doi: 10.1016/j.
conbuildmat.2011.04.032
44. Zhang, Z.; Zhang, B.; and Yan, P., “Comparative Study of Effect
of Raw and Densified Silica Fume in the Paste, Mortar and Concrete,”
Construction and Building Materials, V. 105, 2016, pp. 82-93. doi:
10.1016/j.conbuildmat.2015.12.045
45. Sun, K.; Zhou, X.; Gong, C.; Ding, Y.; and Lu, L., “Influence of Paste
Thickness on Coated Aggregates on Properties of High-Density Sulphoalu-
minate Cement Concrete,” Construction and Building Materials, V. 115,
2016, pp. 125-131. doi: 10.1016/j.conbuildmat.2016.04.018
46. Quanji, Z.; Lomboy, G. R.; and Wang, K., “Influence of Nano-Sized
Highly Purified MagnesiumAlumino Silicate Clay on Thixotropic Behavior
of Fresh Cement Pastes,” Construction and Building Materials, V. 69, 2014,
pp. 295-300. doi: 10.1016/j.conbuildmat.2014.07.050
47. Mechtcherine, V.; Nerella, V. N.; and Kasten, K., “Testing Pumpability
of Concrete Using Sliding Pipe Rheometer,” Construction and Building
Materials, V. 53, 2014, pp. 312-323. doi: 10.1016/j.conbuildmat.2013.11.037
48. Wallevik, O. H.; Feys, D.; Wallevik, J. E.; and Khayat, K. H.,
“Avoiding Inaccurate Interpretations of Rheological Measurements for
Cement-Based Materials,” Cement and Concrete Research, V. 78, 2015,
pp. 100-109. doi: 10.1016/j.cemconres.2015.05.003
49. Koehler, E., and Fowler, D., “Development of a Portable Rheometer
for Fresh Portland Cement Concrete,” ICAR Report:103-105, 2004.
50. Kovler, K., and Roussel, N., “Properties of Fresh and Hardened
Concrete,” Cement and Concrete Research, V. 41, No. 7, 2011, pp. 775-792.
doi: 10.1016/j.cemconres.2011.03.009
51. Xie, H. B.; Liu, F.; Fan, Y. R.; Yang, H. Y.; Chen, J.; Zhang, J.; and
Zuo, C. G., “Workability and Proportion Design of Pumping Concrete
Based on Rheological Parameters,” Construction and Building Materials,
V. 44, 2013, pp. 267-275. doi: 10.1016/j.conbuildmat.2013.02.051
52. Acrivos, A., “Shear-Induced Particle Diffusion in Concentrated
Suspensions of a Noncolloidal Particles,” Journal of Rheology (New York,
N.Y.), V. 39, No. 5, 1995, pp. 813-826. doi: 10.1122/1.550619
53. Roussel, N., and Cussigh, F., “Distinct-Layer Casting of SCC: The
Mechanical Consequences of Thixotropy,” Cement and Concrete Research,
V. 38, No. 5, 2008, pp. 624-632. doi: 10.1016/j.cemconres.2007.09.023
@seismicisolation
@seismicisolation
21. 20 ACI Materials Journal/November 2021
Thereafter, the perceived rheological improvements are
practically captured by performing buildability tests.13
Hardened state mechanical properties, including compres-
sive and flexural strengths and Young’s modulus, are deter-
mined. Additionally, the influence of improved thixotropy
via nanoparticle addition on the interlayer bond strength14
is
investigated via four-point bending tests.
RESEARCH SIGNIFICANCE
The addition of nanoparticles to 3D-printable concrete
materials may improve thixotropy behavior that will conse-
quently aid the general 3D printing constructability process—
namely pumpability, extrudability, filament shape retention,
and buildability. In the case of LWFC, buildability perfor-
mance might be improved to an acceptable degree where
the requirement for chemical accelerators is negated. In the
case of HPC, higher vertical building rates might be achieved
together with constructability optimization15
without requiring
chemical accelerators, which generally is detrimental towards
obtaining mechanically strong interlayers.An added benefit of
nanoparticle addition is the improved hardened state mechan-
ical properties obtained at low dosages.
EXPERIMENTAL PROCEDURE
Three cases are investigated in this research. Initially,
nS particles are added to the HPC by 1, 2, and 3% (cement
mass) dosages. Similarly, SiC particles are added to the HPC
by 1, 2, and 3% cement mass dosages. Lastly, nS particles
are added to LWFC by 2 and 3% cement mass dosages. The
nS particles are mainly used to improve rheology behavior
and hence, 3D printing constructability performance. These
nanoparticles have a specific surface area (SSA) that is 16
times that of the SiC nanoparticles used in this study, due
to them being twice as small in diameter. The SiC nanopar-
ticles, on the other hand, possess greater mechanical prop-
erties and are envisaged to improve hardened concrete state
mechanical properties. This section describes the exper-
imental procedures followed to conduct all the material
preparation and characterization tests.
Materials
The properties of the nS and SiC nanoparticles employed
in this study are given in Table 1. The nanoparticles are used
in powder form, added together with the other dry constit-
uents and mechanically mixed using a two-bladed pan
mixer. Water was added after 2 minutes’ mixing, which then
continued for a further 2 minutes. A modified polycarbox-
ylate ether high-range water-reducing admixture (HRWRA)
is used to disperse the nanoparticles (detailed informa-
tion on relevant dispersion mechanisms can be found in
Reference 4), and added to the material whereafter mixing
continued for a further 2 minutes. Eleven different 3D-
printable concrete mixtures are investigated in this study—
seven HPCs given in Table 2 and four LWFCs given in
Table 3. CEM II/A-L 52.5N cement is used containing
Table 1—Silica and silicon carbide nanoparticles’
material properties8
Silica (nS) Silicon carbide (SiC)
Purity 99.5% 97.5%
Average particle size (APS) 15 to 20 nm 45 to 55 nm
Specific surface area (SSA) 640 m2
/g 35 to 40 m2
/g
Color White Dark gray
Morphology Spherical Spherical
Bulk density 0.1 g/cm3
0.068 g/cm3
True density 2.2 to 2.6 g/cm3
3.22 g/cm3
Synthesis method — Plasma CVD
Table 2—Reference HPC mixture and nanoparticle mixtures’ constituents and quantities
Mixture name
Mixture constituents and quantities, kg
Cement Sand Fly ash Silica fume Water SP nS SiC
HPC_ref 579 1167 165 83 261 12.2 — —
HPC_nS1 573.2 1167 165 83 261 12.2 5.8 —
HPC_nS2 567.4 1167 165 83 261 12.2 11.6 —
HPC_nS3 561.6 1167 165 83 261 12.2 17.4 —
HPC_SiC1 573.2 1167 165 83 261 12.2 — 5.8
HPC_SiC2 567.4 1167 165 83 261 12.2 — 11.6
HPC_SiC3 561.6 1167 165 83 261 12.2 — 17.4
Table 3—Reference LWFC mixture and nanoparticle mixtures’ constituents and quantities
Mixture name
Wet density, kg/m3
Mixture constituents and quantities, kg
Pre–post 3DP Cement Fly ash Water PP Foam nS
LWFC_ref 1410–Not measured 506.2 506.2 369.6 4.1 18 —
LWFC_nS2 1435–1445 506.2 506.2 369.6 4.1 18 11.6
LWFC_nS3 1445–1461 506.2 506.2 369.6 4.1 18 17.4
LWFC_nS2_new*
1648–1720 524.8 524.8 330.6 4.1 19.8 11.6
*
Adjusted mixture to reduce total water content.
@seismicisolation
@seismicisolation
22. 21
ACI Materials Journal/November 2021
between 6 and 20% limestone extender. The SCMs include
Class S fly ash, which is equivalent to Class F according
to ASTM C618,16
together with ferroatlantica silica fume,
mainly used for enhanced rheology properties. The aggregate
is a locally mined fine Malmesbury sand that consists of a
continuously graded particle size distribution and maximum
particle size of 4.75 mm. In all instances, ordinary potable
tap water is used. When preparing the LWFC mixtures,
the nanoparticles are added to the water and mechanically
mixed as a means of dispersion. Polypropylene (PP) fibers, 6
mm in length and 40 μm in diameter, are added to the LWFC
to enhance fresh state stiffness. The LWFC comprises the
fly ash and cement paste base mixed to a fluid consistency,
which is subsequently combined and thoroughly mixed with
precursor foam of density 75 kg/m3
made from a protein-
based foam agent solution with a water-solution dilution
ratio of 1:40 and ferrous sulfate as a stabilizer with a dilution
water-powder ratio of 1:80. With the addition of pre-mixed
and predetermined amounts of foam, the target wet densities
of 1400 ± 50 and 1600 ± 50 kg/m3
investigated in this paper
are achieved. These wet densities relate to dry densities
of roughly 1230 and 1450 kg/m3
, which have been shown
to be applicable for structural use as load-bearing walls in
buildings.17,18
Current work on 3D-printable LWFC of wet
densities between 700 and 1000 kg/m3
for load-bearing
low-rise structural walling is advancing and to be reported
in a future publication.19
Rheology characterization
To capture thixotropy behavior pertaining to 3D printing
of concrete, the bilinear thixotropy model developed by
Kruger et al.4
is employed in this study. The model distin-
guishes between short term re-flocculation (Rthix), which
is a physical, interparticle force equilibrium process, and
long-term structuration (Athix), which is a chemical process
mostly grounded in hydration kinetics. Higher Rthix values
indicate increased thixotropy behavior, whereas higher Athix
values imply reduced settings time of the material. Further-
more, lower dynamic yield shear stresses (τD,i) are favorable
for pumping of a material, whereas higher static yield shear
stresses (τS,i) improves 3D printing buildability. Figure 1
depicts the bilinear static yield shear stress evolution model
employed in this research to quantify material rheology.
This model is developed using data obtained from multiple
stress growth tests, in particular the maximum and post-peak
minimum stresses for each test which corresponds to the
static and dynamic yield stresses respectively, conducted at
various resting time intervals with an ICAR rheometer.4,20
A
shear rate of 1/s is employed for the HPC rheometry testing,
which is the approximate shear induced by the concrete
pump during 3D printing. The LWFC that typically is much
more viscous, is tested at a shear rate of 0.124/s due to the
lower 3D printing speed employed in the buildability test,
thus yielding reduced pump speed and pressure. For both
materials, a stress growth test duration of 60 seconds is
employed. By plotting the recorded τS,i, τD,i, Rthix, and Athix
values, the bilinear static yield shear stress evolution curve
is obtained, as depicted in Fig. 1.
Buildability performance characterization
Buildability in terms of 3D printing of concrete refers to
the number of filament layers that can be deposited onto one
another in a successive manner without excessive deforma-
tion occurring that could result in in-print collapse.13
The
higher the total build height and vertical building rate, the
better the buildability performance. Failure typically occurs
due to either plastic yielding (material) or elastic buckling
(instability).21
To solely focus on material plastic yielding,
long straight elements must be avoided and element geometry
selected such that near equal moment of inertias are obtained
about all axes, in which case a circular hollow column is
widely employed.22
For this research, a 250 mm diameter
circular hollow column is printed at 60 mm/s nozzle speed,
10 mm deposition and layer height obtained from a 25 mm
diameter circular nozzle, yielding a layer width of approx-
imately 35 mm. Due to the LWFC being more flowable
than the HPC, thus expecting significantly less comparative
buildability, a print speed of 40 mm/s is employed to achieve
failure at a higher number of layers. This yields a vertical
building rate of 0.05 m/min for the HPC and 0.03 m/min
for the LWFC. For all buildability tests, a column is printed
until failure occurs and the total number of deposited layers
achieved at failure captured.
Mechanical properties characterization
The mechanical properties of 3D-printable mixtures
are determined through conventional casting and testing
methods. Printed samples are not considered as the inter-
layer bond significantly contributes to the overall strength
observed in various testing directions, which is not ideal for
comparative purposes. The mechanical properties consid-
ered are the compressive and flexural strength, as well as
Young’s modulus characterization. At least three specimens
were tested for each mechanical property and 3D-printable
material combination. Note that, in the case of the LWFC,
only LWFC_ref and LWFC_nS2 are considered (refer to
Table 3) for mechanical testing, due to the limited total
available amount of nanoparticles, justified by the focus on
rheology. Concrete testing ages include 1, 7, 28, and 56 days
for the HPC mixtures and 14 and 28 days for the LWFC.
Altogether, a total of at least 274 specimens were tested. All
specimens are initially cured at laboratory conditions while
still in the molds; however, a day after casting, all specimens
Fig. 1—Bilinear static yield shear stress evolution model
pertainingto3Dprintingofconcreterheologyquantification.4
@seismicisolation
@seismicisolation
23. 22 ACI Materials Journal/November 2021
are demolded and cured in water at 23 ± 2°C until their
respective concrete testing ages are reached.
The flexural and compressive strengths are characterized
according to EN 196-1.23
Flexural specimens with dimen-
sions 40 x 40 x 160 mm are tested in four-point bending to
ensure little to no shear is present where failure is expected
to occur—that is, pure bending behavior is investigated. The
specimens are tested in a material testing machine. A span
of 150 mm is employed with loading points (via rollers)
spaced at 50 mm intervals. A vertical load is transferred to
two rollers and onto the specimen at a rate of 50 N/s. After
conducting the flexural tests, an undamaged 40 x 40 mm
segment of the flexurally tested prism half is tested in
compression.23
Platens of 40 x 40 x 10 mm are placed above
and below the specimen for accurate load transfer onto the
specimen. Specimens are uniaxially loaded at a rate of 2400
N/s. The Young’s modulus is determined according toASTM
C469-02.24
Cylindrical test specimens of 100 mm diam-
eter and 200 mm height are circumferentially instrumented
with three linear variable differential transducers (LVDTs)
at 120-degree intervals, over a 70 mm gauge length. The
testing is conducted in a 2 MN press machine with a 2 MN
loadcell and measurements captured via an data acquisition
system. The specimens are uniaxially loaded up to 40% of
their ultimate compressive capacity for the Young’s modulus
determination.
Interlayer bond strength characterization
The interlayer bond strength (IBS) is determined from
3D-printed specimen loaded in four-point bending23
with
filament layers oriented vertically (O-III),14
as depicted
in Fig. 2. The IBS specimens with dimensions 40 x 40 x
160 mm are extracted via saw cutting from a 175 x 175 x
350 mm square hollow column. This column is printed at a
nozzle speed of 60 mm/s, yielding layer heights and widths
of 10 and 40 mm, respectively. Altogether, a pass time
(that is, the time between successive depositions of fila-
ment layers) of 12 seconds is obtained.25
Only the 28-day
concrete testing age is considered in this case. Furthermore,
only the mixtures that yielded the best improvement in terms
of thixotropy are considered due to the limited amount of
nanoparticles available. The reduction in bond strength is
determined using the following equation14
SR (%) = (1 – (SIII/Sm)) ∙ 100 (1)
with SR the effective IBS reduction in % compared to the
mold-cast specimens (Sm); and SIII the 3D-printed elements’
flexural strengths determined in the test series in O-III.
EXPERIMENTAL RESULTS AND DISCUSSION
High-performance 3D-printable concrete
Static yield shear stress evolution and thixotropy—
The static yield shear stress evolution curves for the HPC
mixtures containing nS and SiC nanoparticles are depicted
in Fig. 3 and the respective model parameters given in
Fig. 2—3D-printed 175 x 175 x 350 mm square hollow column from which 40 x 40 x 160 mm IBS specimens are extracted and
tested in four-point bending according to EN 196-1.23
Fig. 3—Bilinear static yield shear stress evolution curves
for: (a) nS HPC; and (b) SiC HPC 3D printable mixtures.
@seismicisolation
@seismicisolation
24. 23
ACI Materials Journal/November 2021
Table 4. The same HPC reference mixture was used for the
SiC investigations as for the nS investigations; however,
discrepancies are observed between the two listed reference
HPC rheology properties. This is due to the SiC investiga-
tions being conducted at a different time period than that of
the nS, and fresher constituent materials were used for the
SiC investigations. A drastic increase in static yield stress is
observed for the nS addition for every percentage increment.
For 2% nS addition, a 100% increase in static yield stress is
observed from the reference mixture. However, it was visu-
ally observed that the HPC_nS2 and HPC_nS3 mixtures do
not possess the required workability to be pumpable. One
percent nS addition was found to increase the thixotropy,
specifically Rthix, by 16%. However, Rthix deteriorated for
every percentage nS addition thereafter. Athix reduced for
every percentage increment of nS addition. The SiC nanopar-
ticles significantly improved the rheology behavior of the
HPC_ref*
mixture. In particular, a 149 and 141% improve-
ment in Rthix and Athix was observed, respectively, for 1%
SiC addition. However, only a maximum of 30% increase in
static yield stress is obtained with SiC addition, compared to
137% with nS addition. This is mainly ascribed to the SSA
of the nanoparticles, of which the nS’s SSA is more than
16 greater than that of the SiC nanoparticles. Consequently,
larger water or HRWRA dosage is required to disperse the
(nano) particles and release entrapped water. Hence the
drastic increase in static yield stress observed by the nS
addition. In the case of the SiC presenting less SSA, better
dispersion of particles is obtained, which therefore does
not significantly influence static yield stress. Instead, the
tremendous number of dispersed nanoparticles contribute to
the thixotropy mechanism, thus especially improving Rthix.
Buildability performance—The circular hollow columns
printed to assess the buildability performance of the HPC
with nanoparticle inclusions are depicted in Fig. 4. The
optimum nS addition for enhanced thixotropy and thus
3DCP constructability was found to be 1%, based on the
rheology results given in the previous section. For SiC, the
optimum dosage was determined as 2%, because not only
is high thixotropy important for 3DCP, but also appropriate
static yield stress. Therefore, HPC_nS1 and HPC_SiC2 were
selected for the buildability validation prints and compared
to their respective reference mixtures, namely HPC_ref
and HPC_ref*
. Only these two nanoparticle mixtures were
considered due to the limited amount of particles available.
The HPC_ref mixture attained a total build height of
540 mm (54 layers) before failure occurred, which is 13
layers less than that of the HPC_ref*
mixture (670 mm).
Based on the bilinear yield stress data, it was expected that
the HPC_ref mixture would perform better than the HPC_ref*
mixture in the buildability test. The reason for this discrep-
ancy is unknown. The HPC_nS1 mixture attained a total
build height of 590 mm—that is, five filaments layers more
than the HPC_ref mixture. This result was expected based
on the improved rheology properties enabled by the 1% nS
addition to the reference mixture. The HPC_SiC2 attained a
Table 4—HPC mixtures’ yield stresses and thixotropy parameters
HPC ref HPC nS1 HPC nS2 HPC nS3 HPC ref*
HPC SiC1 HPC SiC2 HPC SiC3
τS,i, Pa 2730 3944 5618 6483 1545 1647 2007 1636
τD,i, Pa 1146 1532 2702 2803 560 480 738 966
Rthix, Pa/s 6.88 8 6.11 4.2 2.04 5.07 3.9 3.01
Athix, Pa/s 1.08 0.91 0.61 0.61 0.32 0.77 0.63 0.66
*
SiC investigations were performed at different time than that of nS investigations. Subsequently, this HPC reference mixture exhibited dissimilar rheological characteristics than
exact HPC reference mixture employed at time of nS investigations.
Fig. 4—Buildability performance quantification via printing
circular hollow column for: (a) HPC_ref versus; (b) HPC_
nS1; and (c) HPC_ref*
versus (d) HPC_SiC2.
@seismicisolation
@seismicisolation
