Guide to Durable
Concrete
ACI 201.2R-16
Reported by ACI Committee 201
First Printing
November 2016
ISBN: 978-1-945487-39-2
Guide to Durable Concrete
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ACI 201.2R-1 6
Guide to Durable Concrete
Reported by ACI Committee 201
Thomas J. Van Dam, Chair
Reza Ahrabli
James M. Aldred
Jon B. Ardahl
Mohamed Bassuoni
Bruce Blair
Andrew J. Boyd
Paul W. Brown
Ramon L. Carrasquillo
Rachel J. Detwiler
Jonathan E. Dongell
R. Douglas Hooton, Secretary
Thano Drimalas
Kevin J. Folliard
Harvey H. Haynes
Jason H. Ideker
Francis Innis
Donald J. Janssen
Roy H. Keck
Mohammad S. Khan
Kimberly E. Kurtis
Michael L. Leming
Tyler Ley
Darmawan Ludirdja
Mohamad Nagi
Robert E. Neal
Charles K. Nmai
Karthik H. Obla
Robert C. O’ Neill
Kyle Austin Riding
David A. Rothstein
Hannah C. Schell
Lawrence L. Sutter
David G. Tepke
Michael D. A. Thomas
Paul J. Tikalsky
David Trejo
Orville R. Werner II
Terry J. Willems
Michelle L. Wilson
Consulting Members
W. Barry Butler
Bernard Erlin
Odd E. Gjorv *
William G. Hime
*
Charles J. Hookham
Alexander M. Leshchinsky
Stella Lucie Marusin
Howard H. Newlon Jr.
Mauro J. Scali
George V. Teodoru
Niels Thaulow
J. Derle Thorpe
Claude B. Trusty Jr.
Deceased.
This guide describes specifc types o f concrete deterioration. Each
chapter contains a discussion o f the mechanisms involved and the
recommended requirements for individual components o f concrete,
quality considerations for concrete mixtures, construction proce dures, and in fuences o f the exposure environment, which are all
important considerations to ensure concrete durability.
This guide was developed for conventional concrete but is generally applicable to specialty concretes; however, specialty concretes,
such as roller-compacted or pervious concrete, may have unique
durability-related issues that deserve further attention that are not
addressed herein.
Keywords: abrasion resistance; alkali-aggregate reaction; chemical attack;
curing; deterioration; durability; freezing and thawing; physical salt attack,
sulfate attack.
ACI Committee Reports, Guides, and Commentaries are
intended for guidance in planning, designing, executing, and
inspecting construction. This document is intended for the use
o f individuals who are competent to evaluate the signifcance
and limitations of its content and recommendations and who
will accept responsibility for the application of the material it
contains. The American Concrete Institute disclaims any and
all responsibility for the stated principles. The Institute shall
not be liable for any loss or damage arising therefrom.
Reference to this document shall not be made in contract
documents. If items found in this document are desired by
the Architect/Engineer to be a part of the contract documents,
they shall be restated in mandatory language for incorporation
by the Architect/Engineer.
CONTENTS
CHAPTER 1 —INTRODUCTION AND SCOPE, p. 2
1 .1 —Introduction, p. 2
1 .2—Scope, p. 3
CHAPTER 2—DEFINITIONS, p. 3
2.1 —Defnitions, p. 3
CHAPTER 3—MASS TRANSPORT, p. 3
—
—
—
3.1 Introduction, p. 3
3.2—Transport processes in nonreactive porous media, p. 4
3.3 Factors affecting mass transport in concrete, p. 5
3.4—Measurement of transport properties, p. 8
3.5 Obtaining durable concrete, p. 1 0
CHAPTER 4—FREEZING AND THAWING OF
CONCRETE, p. 10
4.1 —Introduction, p. 1 0
ACI 201 .2R-1 6 supersedes ACI 201 .2R-08 and was adopted and published
November 201 6
Copyright © 201 6, American Concrete Institute.
All rights reserved including rights of reproduction and use in any form or by
any means, including the making of copies by any photo process, or by electronic
or mechanical device, printed, written, or oral, or recording for sound or visual
reproduction or for use in any knowledge or retrieval system or device, unless
permission in writing is obtained from the copyright proprietors.
2
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
4.2—Frost attack of concrete made with durable aggregates, p. 11
4.3—Frost attack of concrete made with nondurable
aggregates, p. 1 7
CHAPTER 5—ALKALI-AGGREGATE REACTION,
p. 1 9
5.1 —Introduction, p. 1 9
5.2—Types of reactions, p. 1 9
5.3—Evaluating aggregates for potential alkali-aggregate
reactivity, p. 22
5.4—Preventive measures, p. 25
5.5—Tests for evaluating preventive measures, p. 28
5.6—Protocols for minimizing the risk of alkali-aggregate
reactivity, p. 29
CHAPTER 6—SULFATE ATTACK, p. 30
6.1 —External sulfate attack, p. 30
6.2—Internal sulfate attack, p. 36
6.3—Seawater and brine exposure, p. 37
CHAPTER 7—CHEMICAL ATTACK, p. 39
7.1 —General, p. 39
7.2—Seawater, p. 39
7.3—Acid attack, p. 41
7.4—Fresh water, p. 42
7.5—Carbonation, p. 42
7.6—Industrial chemicals, p. 43
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7.7—Deicing and anti-icing
chemicals,
44
7.8—Environmental structures, p. 45
from
CHAPTER 8—PHYSICAL SALT ATTACK, p. 45
8.1 —Introduction, p. 45
8.2—Occurrence, p. 46
8.3—Background, p. 47
8.4—Mechanism, p. 47
8.5—Recommendations, p. 48
CHAPTER 9—CORROSION OF METALS
AND DEGRADATION OF OTHER MATERIALS
EMBEDDED IN CONCRETE, p. 48
9.1 —Introduction, p. 48
9.2—General principles of corrosion initiation in concrete,
p. 48
9.3—Propagation of corrosion, p. 49
9.4—Corrosion-related properties of concreting materials,
p. 49
9.5—Mitigating corrosion, p. 50
9.6—Corrosion of prestressed steel reinforcement, p. 53
9.7—Degradation of materials other than steel, p. 53
9.8—Summary, p. 54
CHAPTER 10—ABRASION, p. 54
1 0.1 —Introduction, p. 54
1 0.2—Testing concrete for resistance to abrasion, p. 55
1 0.3 —Factors affecting abrasion resistance of concrete,
p. 55
1 0.4—Recommendations for obtaining abrasion-resistant
concrete surfaces, p. 57
1 0.5—Studded tire and tire chain wear on concrete, p. 58
1 0.6—Skid resistance of pavements, p. 58
1 0.7—Erosion, p. 59
CHAPTER 11 —SUMMARY, p. 60
CHAPTER 1 2—REFERENCES, p. 60
Authored documents, p. 62
CHAPTER 1 —INTRODUCTION AND SCOPE
1.1 —Introduction
Concrete is the most widely used construction material in
the world. The design, detailing, and execution of concrete
to resist weathering action, chemical attack, abrasion, and
other processes of deterioration over its intended service life
will determine its durability. Durable concrete will retain its
original form, quality, and serviceability when exposed to its
environment. Properly designed, proportioned, transported,
placed, fnished, and cured concrete is capable o f providing
decades of service with little or no maintenance. Yet certain
conditions or environments exist that can lead to concrete
deterioration. Deterioration mechanisms are either chemical or physical in nature and may originate from within the
concrete, or may be the result of the external environmental
exposure. Chemical and physical attacking mechanisms often
Standard
Sharing Group
and our
chats
work synergistically.
Depending
on the
nature of the attack,
distress may be concentrated in the paste, aggregate, or reinforcing components of the concrete, or a combination thereof.
The various factors in f uencing durability and a particular
mechanism of deterioration should be considered in the
context of the environmental exposure of the concrete. In
addition, consideration should be given to the microclimate
to
which
the
s pecifc
s tructural
element
is
to
be
expos ed.
The type and severity of deterioration of a given structure
may be affected by its proximity to sources of deleterious
agents or agents that facilitate distress, exposure to wind,
precipitation, or temperature. For instance, exterior girders
in a bridge structure may be exposed to a more aggressive
environment than interior girders.
The concept of service life is increasingly used for the
design of new structures. To produce concrete suitable for a
particular application, required service life, design requirements, and expected exposure environments, both macro
and micro, should be determined be fore defning the neces sary materials and mixture proportions.
The use of good materials and proper mixture proportioning
will not, by itself, ensure durable concrete. Appropriate placement practices and workmanship are essential to the production of durable concrete. Fresh concrete can be consolidated
and molded to the shape desired to serve its intended purpose.
During this stage, a number o f properties signifcantly inf uencing the durability of the hardened concrete are established.
Pore structure development, air-void system formation, material uniformity, and potential for cracking are established
at early ages and are important to the ultimate durability of
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
concrete. As such, durable concrete requires the application
of good quality control during construction. Inspection and
testing by trained and certifed personnel
can help ensure the
use of durable mixtures and proper practices.
1.2—Scope
This guide discusses the important mechanisms of co ncrete
deterioration and gives recommendations on how to mitigate
or minimize such damage. This guide also addresses durability by frs t dis cuss ing the importance o f mass transport
and then addres sing
s pecifc
modes
o f attack in s eparate
chapters. These include freezing and thawing, alkali-aggregate reaction (AAR), sulfate attack, aggressive chemical
attack, physical salt attack, corrosion of metals and other
embedded materials, abrasion, or a combination of these.
Fire resistance of concrete and cracking are not addressed
directly. Fire resistance is covered in ACI 21 6.1 and cracking
is covered in ACI 224R and ACI 224.1 R. While cracking
does impact the durability of concrete in severe exposures,
the di fferent causes o f cracking and their s pecifc impacts are
not discussed. Cracking is only mentioned in general terms
regarding its impact on
f
uid ingres s .
CHAPTER 2—DEFINITIONS
2.1 —Defnitions
ACI provides a comprehensive lis t o f defnitions through
an online resource, “ACI Concrete Terminology,” https://
www.concrete.org/store/productdetail.aspx?ItemID=CT 1 6.
D efnitions
provided herein complement that s ource.
advective transport ––transfer of heat or matter via the
bulk motion o f a f uid.
alkali loading (or content) ––total amount of equivalent
alkalis (Na 2 O e) in a concrete mixture expressed as mass per
volume.
calcium sulfoaluminate cement ––product obtained
by
pulverizing
clinker
containing
mainly
ye′elimite
[Ca4 (AlO 2 ) 6 SO 4 ] that is often used in expansive cements and
ultra-high-early-strength cements.
diffusion ––movement of species, such as ions, gas, or
vapor, from an area of higher concentration to an area of lower
concentration, independent o f the bulk motion o f a f uid.
electrical migration ––transport of electrons or ions due
to an electric potential gradient.
ice lens ––layer of ice, generally parallel to the exposed
surface of the concrete, that can produce internal damage
and also lead to scaling or delamination.
leaching ––dissolution and removal of soluble components such as calcium hydroxide from concrete.
permeability —the ability of a given concrete to permit
liquids or gases to pass through.
permeation ––f ow o f a liquid, gas , or vapor within a solid
under the action of a pressure gradient.
physical salt attack ––mechanism in which concrete or
mortar is damaged as a result of salt crystallization pressure.
reactive silica ––form of silica, often amorphous or
crypto-crystalline, that dissolves when in contact with
3
p o re s o l uti o n havi ng a s u ff c i e ntl y hi g h c o nc e n tration of hydroxyl ions.
salt weathering ––form of deterioration most commonly
observed in arid climates where exposure to soluble salts
and cyclic variations in temperature and relative humidity
can lead to salt crystallization.
thaumasite ––silicate mineral, colorless to white prismatic hexagonal crystals typically as acicular radiating
groups, with the chemical formula {[Ca 3 Si(OH) 6 · 1 2(H 2 O)]
(SO 4 )(CO 3 )} .
c o nc re te
CHAPTER 3—MASS TRANSPORT
3.1 — Introduction
Concrete is a multiphase porous medium consisting of a
multiscale porous cement paste matrix with aggregate inclusions. Liquid and gas may be present in any pores and microcracks. As such, it is susceptible to the ingress and movement o f s ubstances
( f uids or ions) from its environment
within and through its pore system. This chapter discusses
the transport of gases, liquids, and ions in solution through
concrete (Lichtner et al. 1 996; Baer 1 988; Hearn et al. 2006;
Hall and Hoff 201 2). Methods for improving the durability
of concrete and some of the common test methods used to
measure the transport properties, along with their advantages
and limitations with regard to assessing concrete durability,
are also discussed. It is recognized that the rate of ingress
o f f uids and ions will increas e by the pres ence o f cracks.
However, the specifc in f uences o f di fferent types o f cracks
and crack widths are not discussed herein.
The ingress of gases, liquids, or ions in solution through
concrete may initiate chemical processes, physical processes,
or both, that affect the durability of the concrete under a given
set of service conditions. Water itself may be harmful because
of its ability to leach calcium hydroxide (CH) from the hardened cement paste and because of osmotic pressures generated as water f ows to sites o f higher alkalinity ( Powers et al.
1 954; Powers 1 975; Helmuth 1 960b,c). In addition, water
may also be acidic or carry harmful dissolved chemicals, such
as chlorides or sulfates, into the concrete. The ingress of gases
such as oxygen and carbon dioxide through the concrete pores
can contribute to the corrosion of steel reinforcement.
Different substances may interact with components of the
concrete in different ways; therefore, transport of a substance
through concrete is unique to that substance. For example,
water can hydrate previously unhydrated cement particles or
leach calcium. Chloride ions may be bound by the hydration
products of cement or supplementary cementitious materials
(SCMs). The size of the molecules or ions that are transported through the concrete, vis cosity o f the f uid, valence o f
the ions, and other ionic species present also affect the transport properties. Thus, permeability and diffusivity must be
expressed in terms of the substance that is migrating through
the concrete. In general, concrete with transport properties
that limit the rate of ingress of external agents is not immune
to chemical deterioration, but the effects are mainly near the
exposed surfaces, so the concrete tends to be more durable.
4
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
3.2—Transport processes in nonreactive porous
media
This section provides a brief overview of the transport
o f f uids (gases and liquids ) and ions in s olution
within a nonreactive porous medium. This is a simplifying
assumption because concrete changes chemically and physically with time in response to its environment. Physical
changes, chemical changes, or both, in the internal structure
of the porous medium resulting from interactions with the
migrating f uids or ions are not dis cus s ed herein. The trans port pathways described here include:
a) Transport by permeation
b) Advective transport
c) Hydrodynamic dispersion
d) Diffusion within the pores
e) Transport due to electrostatic interactions or electrical
migration
Martín-Pérez et al. (2001 ) modeled transport related to
corrosion of reinforcement in concrete based on chloride
transport, moisture diffusion, heat transfer, and oxygen
proces ses
transport
us ing
a
two- dimensional
fnite
element
model.
f F ick’s s econd law.
Johannesson (2003) developed a theoretical model for
diffusion of different types of ions in concrete pore solution.
The model incorporates diffusion caused by concentration
gradients of ions (for example, due to drying), internal electrical potential, convection, effects of changes in moisture
content, and mass exchange of ions between solution and
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Chung and Consolazio (2005) developed a fnite difference
model to simulate heat and mass transport in rapid heating
conditions, such as fres in rein forced concrete structures. The
model accounts for the interference between liquid and gas
phases, slip- f ow e ffects in steam f ow, and the inter ference o f
steel reinforcement in moisture movement in concrete.
3.2.1 Transport by permeation —Permeation is the f ow o f
a f uid under the action o f a press ure gradient. Permeability
They used a modifed version o
is
the
property
that
characterizes
the
eas e
with
which
f
uid
passes through a porous material under a pressure gradient.
F or
a
s teady
medium,
the
laminar
f
uid
f
ow
f
ow
is
through
related
to
a
the
s aturated
hydraulic
porous
pres s ure
gradient according to Darcy’s law.
dq /dt = K1 A h /l
(3.2.1 )
∆
where
dq /dt
ability
coe
K
is the f ow (expres s ed as a rate) ;
1 is the permeffcient;
is the cros s - s ectional area; ∆
is the
hydraulic head; and l is the thickness of the specimen. The
permeability coe ffcient
1 is the rate of discharge of water
under laminar
A
f
h
K
ow conditions through a unit cros s - sectional
area of a porous medium under a unit hydraulic gradient
and standard temperature conditions. Darcy’s law indicates
that for a given cross - s ectional area and permeability coe ff cient, the f ow is proportional to the hydraulic gradient ∆ /l.
Under s ervice conditions ,
f
h
ow is three- dimens ional
and the
concrete may not be saturated. In concrete, the permeability
ffcient may change with increas ed hydration, cracking,
coe
Fig. 3.2.1‒Graphical representation o f a simple percolation
theory model used to model the permeability o f concrete.
or changes in the pore structure due to various physical and
chemical processes.
Permeability coe ffcients o f plastic portland cement pastes
of 0.5 water-cement ratio (w/ ), calculated from measurements
of bleeding, ranged from 5 to 8 × 1 0 –7 m/s for four cements
with di fferent chemical composition but the same specifc
surface (1 80 m2 /kg by the Wagner turbidimeter). The permeability coe ffcient o f mature paste ( for example, at greater than
28 days Sharing
of age) is between
millionth
and one 1 0 millionth of
Standard
Group 1and
our chats
that of fresh paste. It ranges from 1 × 1 0 –1 5 to 1 .2 × 1 0 –1 2 m/s
for w/ ranging from 0.3 to 0.7 (Powers et al. 1 954).
To understand the effects of microstructure on the permeability of concrete, Bentz et al. (1 999) at the National Institute of Standards and Technology (NIST) used percolation
theory. One useful application of percolation theory is the
examination of the time needed for a material to progress
through a complex maze (Stauffer and Aharony 1 992). This
maze consists of areas that can allow free movement, as
well as areas that impede the trans port o f a f uid to di fferent
degrees. These models can be made in two and three dimensions and can include the effects of cracks. Lu et al. (201 2)
were able to use a three-dimensional version of the NIST
model to predict chloride ingress into cracked concrete.
Work has also been done by NIST that models changes
in the properties of these systems with time. This allows the
change in the microstructure of the concrete to be examined
with time and obs erve the e ffects on the ability o f a f uid to
move through the system.
A graphical representation of a simple percolation
theory model has been used by Bentz (2000) to model
the trans p o rt o f a f ui d thro ug h c o nc re te i n two di me n sions, as shown in Fig. 3 . 2. 1 . Very dense areas are used
for aggregates. Moderately dense material is used for the
cement paste and low-density material is used to model
the interfacial transition zone (3 . 3 . 4).
3.2.2 Advective transport —Advective transport refers to
the movement of molecules and ions with the bulk solution
c
c
f
ow. This trans port proces s is related directly to the velocity
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
5
Fig. 3.3.1a–‒Relative sizes o f different types o f pores and other microstructural features
(adapted from Mehta [1986]).
f fuid fow and concentration o f ions in solution. Depending
on exposure condition and degree of saturation, the velocity
o f f uid f ow depends on the sorptivity or permeability o f
concrete, viscosity o f the f uid phase, and pressure head o f
the permeating f uid or rate o f evaporation from the exposed
surface in the case of wick action (Buenfeld et al. 1 995).
3.2.3 Hydrodynamic dispersion —Dispersion is the spreading
of the ion concentration during advective transport due to
o
variations
in the pore
f
uid velocity.
Thes e variations
dt = D
e
·
2
d c/
dx2
Electrical
migration
occurs
when
an
external
electric
feld
such as in ASTM C1 202 (or AASHTO T277) is applied to
the medium (B uen feld et al. 1 9 9 8 ) . The migration f ux Ji of
ion i is given by
Ji = − Di Ci
zi F ∂ ϕ
RT ∂ x
(3.2.5)
can be
a result of the tortuosity of the pore structure, the connectivity o f the pore network, or variation in the f uid properties .
3.2.4 Diffusion within pores —Diffusion refers to the transport mechanism whereby ions or gases migrate from areas of
higher concentration to areas of lower concentration. For the
idealized one-dimensional case, Fick’s second law describes
the non-steady-state diffusion of ions within the pores
dc /
strated by Snyder (2001 ) and Snyder and Marchand (2001 ).
(3.2.4)
where c is the concentration of the ion at distance x from the
surface after time t, and De is the e ffective di ffusion coe ffcient or effective diffusivity. The effective diffusivity is a
function of the porosity and tortuosity of the porous medium
and the molecular diffusivity of the ion of concern. Many
factors affect diffusion of ions in concrete. Based on measurements obtained under controlled conditions in the laboratory,
di ffusion coe ffcients increase with temperature and waterto-cementitious materials ratio (w/cm ) and decrease with
increasing degree of hydration. Because concrete pore solutions have high ionic strength, electrical charge effects can
be signifcant. Di ffusion coe ffcients can also vary with the
species of other ions present in solution. For these and other
reasons discussed in 3.4.1 .2, the values obtained experimentally using Fick’s second law are generally termed “apparent
di ffusion
coe ffcients”.
Di ffusion
coe ffcients
for Na+ in
–11
–1 3
2
concrete are on the order of 1 0 to 1 0 m /s, and for Cl –, on
the order of 1 0 –11 to 1 0 –1 2 m2 /s (Taylor 1 997).
3.2.5 Transport due to electrostatic interactions or elec trical migration —Migration refers to the transport mechanism due to the charged nature of ions and is the result of
the potential difference across the specimen. The electrical
coupling between ions in concentrated solutions was demon-
x is the potential difference; Zi is the charge of the
ion; F is the Faraday constant; T is the temperature; D i is the
ion diffusivity; R is the gas constant; and Ci is the concentration of the ion in solution.
Further information on ionic transport in concrete can
be found in McGrath and Hooton (1 996) and for the more
complex case of multi-species transport in Truc et al. (2000)
and Samson et al. (1 999).
where ∂φ/∂
3.3 — Factors a ffecting mass transport in concrete
3.3.1
Porosity and pore size distribution —Porosity
f voids as a fraction
expressed as a percent of the total volume
defned
as
the
volume
o
that is
is
usually
porosity (%) = (volume of voids/total volume) × 1 00% (3.3.1 )
Figure 3.3.1 a shows the size ranges for the various types
of pores in concrete. Pores in concrete range in size from
nanometers to millimeters. The capillary pores, also ranging
in size from tens of nanometers to millimeters, have the
mos t signifcant e ffect on the trans port properties .
Trans port properties, however, depend more on the connectivity
of the pores than on either the porosity or size of the pores.
Figure 3.3.1 b shows two hypothetical porous materials with
approximately the same porosity. In one material, the pores
are discontinuous, as would be the case with entrained air
bubbles, whereas in the other the pores are continuous. The
latter material would allow for more rapid rates of transport
than the former.
3.3.2 Water-cement ratio (w/c)—The initial porosity of a
cement paste is determined by the w/c . As cement hydrates,
hydration products fll s ome o f the void s pace formerly occu pied by water. With time, this process results in a continued
6
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
Fig. 3.3.1b–‒Porosity and permeability are related but
distinct. The two hypothetical materials shown have approxi mately the same porosity (total volume o fpores), but different
permeabilities. Discrete pores, such as those resulting from
air entrainment, have almost no effect on permeability, but
interconnected pores increase permeability.
decrease in the porosity of the cement paste. Figure 3.3.2a
(Mehta 1 986) illustrates the relationship among w/c , degree
of hydration, and capillary porosity. For a w/c of 0.45, the
degree of hydration must reach approximately 70 percent to
bring the porosity down to 30 percent. For a w/c of 0.60, the
degree of hydration must reach approximately 1 00 percent
to reach the same porosity. The degree of hydration that
could be expected for good curing conditions—for example,
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moist curing for 5 to 7 days—would range between 70 and
80
percent,
depending
on
cement
chemis try
and
Fig. 3.3.2a‒–Water-cement ratio versus capillary porosity
for cement paste at different degrees o f hydration (Mehta
1986) based on equations developed by Powers and Brown yard (1948).
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fnenes s,
and on hydration temperatures; a 1 00 percent degree of
hydration is not a practical possibility. Figure 3.3.2b shows
the relationship between porosity and permeability. Above
a poros ity o f approximately 3 0 percent, the coe ffcient o f
permeability increases sharply.
Powers (1 962b) calculated that for cement paste with a
w/c o f 0. 3 8 , all o f the capillary pore space was j ust flled by
maximum density gel when all of the cement was hydrated.
Sealed, fully hydrated cement pastes made at w/c above
0.38 have remaining capillary pore space equal to the excess
above 0.38. Partially hydrated mixtures have proportionately less gel and more capillary space. Powers et al. (1 959)
calculated the time required for capillary pores to become
discontinuous with increasing hydration of the cement, as
shown in Table 3.3.2. It is notable that mixtures with a w/c
greater than 0.7 will always have continuous pores. Even for
w/c of 0.40 to 0.45, extended moist curing or other favorable
curing conditions are necessary to achieve the desired low
permeability.
3.3.3 Curing temperature
3.3.3.1 At normal temperatures —Soon after mixing
cement with water, a gel layer forms on the surfaces of the
cement grains (Taylor 1 997). Between 3 and 24 hours after
mixing cement with water, approximately 30 percent of the
cement reacts. Rapid formation of calcium silicate hydrate
(C - S - H) and C H is accompanied by signifcant evolution o f
heat. The CH forms massive crystals in the originally waterflled s pace.
The C- S - H forms a thickening
layer around
Fig. 3.3.2b‒–Both compressive strength and permeability
are related to the capillary porosity o f the cement paste
(adapted from Powers [1958]).
Table 3.3.2–Approximate age required to
produce maturity at which capillaries become
discontinuous for concrete continuously moistcured (Powers et al. 1 959)
w/c by mass
Time required
0.40
3 days
0.45
7 days
0.50
1 4 days
0.60
6 months
0.70
1 year
Over 0.70
Impossible
the cement grains. As the shells grow outward, they begin
to coalesce at about 1 2 hours, a time coinciding with the
maximum rate of heat evolution (Fig. 3.3.3.1 ) and corresponding approximately to completion of setting (Taylor
1 997).
In
F ig.
3.3.3.1 ,
the
frs t
heat
peak
is
as s ociated
with the initial hydrolysis of the C 3 S and the hydration of
the C 3 A. The acceleration period begins with the renewed
evolution of heat at the beginning of the second peak as the
initial hydration products of the C 3 S begin to form. Initial
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
7
Fig. 3.3.3.1––Heat evolution o f Type I/II portland cement
paste as measured by conduction calorimetry (Image courtesy o f E. Shkolnik.).
set coincides with the beginning of the acceleration period.
Final set takes place just before the maximum point of the
second peak. The shoulder of the second peak is associated with the renewed formation of ettringite (Taylor 1 997).
Further hydration of the cement continues at a much slower
rate, asymptotically approaching 1 00 percent (Mindess and
Young 1 981 ). After the spaces between the hydration shells
and the cement grains fll with hydration products , further
hydration is slow (Taylor 1 997). Capillary pores remaining
in mature cement paste increase in size with w/c and have
diameters of 1 0 nm and higher (Mindess and Young 1 981 ).
3.3.3.2 At high or low temperatures —Like most chemical reactions, cement hydration is faster with increasing
temperature. Verbeck and Helmuth (1 969) postulated that at
elevated temperatures, cement hydration products would not
have time to di ffus e any s ignifcant dis tance from the cement
grain, thus forming relatively dense hydration shells around
the cement grains. A consequence of the uneven distribution
of the solid phases is a coarser pore structure. Goto and Roy
(1 981 ) found that the total porosities of pastes hydrated at
1 40°F (60°C) were greater than those of pastes hydrated at
81 °F (27°C). For cement pastes hydrated at low temperatures, on the order of 50°F (1 0°C), the hydration products
are more evenly distributed and the pores fne and dis con tinuous. For cement pastes hydrated at elevated temperatures, pores are coarser and more interconnected (Kjellsen
et al. 1 991 ) . Cement pas tes containing f y as h, s lag cement,
or both, are less sensitive to the effects of elevated temperatures, as discussed in 3.3.5.
3.3.4 Aggregates —Aggregates generally have fundamentally different transport properties from those of cement
paste. For example, the permeability of granite is typically
two to three orders of magnitude lower than that of cement
paste. The presence of the aggregate in a cement-paste
matrix creates an inhomogeneity in the structure of hardened
concrete known as the interfacial transition zone between
the cement paste and aggregate. Mehta (1 986) reported that,
compared to the bulk cement paste, the interfacial transition
zone has a higher void content, higher contents of CH and
ettringite, reduced content of C-S-H, and larger crystals of
Fig. 3.3.4‒–Representation o f transition zone at paste/
aggregate interface in concrete, showing more coarsely
crystalline and porous microstructure than in interzonal
mass (Mehta 1986).
CH strongly oriented parallel to the aggregate surface (Fig.
3.3.4). Factors contributing to the anomalous nature of the
interfacial transition zone include bleeding, which creates
pockets o f water- flled s pace beneath aggregate
particles ;
les s e ffcient packing o f particles o f cementitious materials
in the vicinity of a surface, which is called the wall effect;
and the one-sided growth effect of dissolved cementitious
materials and hydration products diffusing in from the bulk
cement paste, but not from the aggregate (Bentz et al. 1 995 ).
As the cementitious materials hydrate, the interfacial transition zone flls pre ferentially with CH and ettringite. B ecaus e
of the relatively open space, the crystals can grow large.
Thus, in most concrete, the interfacial transition zone is the
weakest link in terms of mechanical behavior and transport
properties. For the latter, the interfacial transition zone can
serve as a relatively open channel for f uids and ions , and
the CH is vulnerable to leaching and acid attack. For a given
w/c and degree of hydration, water permeability of concrete
made with low-permeability aggregates is approximately
one to two orders of magnitude lower than that of cement
paste due to the interfacial transition zone between aggregate and cement paste (Mehta 1 986). It was found that the
diffusivity of chloride in the interfacial transition zone is 1 0
times greater than that in bulk cement paste (Delagrave et al.
1 997). The connectivity of pores in the interfacial transition
zone may be high, leading to s ignifcantly greater rates o f
transport for some concretes than might be predicted from
their mixture proportions. For mixtures with high coarse
8
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
aggregate
contents ,
such
as
paving
concretes,
f
ow through
the interfacial transition zone can completely dominate
moisture movement in the concrete over a wide range of w/c
(Janssen and Snyder 1 994).
3.3.5 Supplementary cementitious materials —The use of
S CMs can signifcantly reduce the permeability and di ffusivity of concrete. These materials may not reduce the total
poros ity
to
any
great
extent,
but
ins tead
act
to
refne
and
subdivide the pores so that they become less continuous.
S ome benefts are obtained by the improvements in work ability, particularly the reduction in bleeding, afforded by
thes e materials . The us e o f f y as h can als o reduce the water
demand of some concrete mixtures, allowing reductions in
water-to-cementitious materials ratio ( w/cm ) while maintaining
equivalent
workability.
The
lower
specifc
gravity
of SCMs relative to cement results in an increased volume
of solids for a given w/cm with increasing dosage. The
greates t benefts from the standpoint o f durability, however,
derive from the pozzolanic reaction associated with many
SCMs. In this reaction, CH from the hydration of the cement
reacts with noncrystalline silica in SCMs and water to form
C-S-H. Because C-S-H has a greater volume than the CH
and pozzolan from which it forms, the pozzolanic reaction
res ults in a fner s ystem o f capillary pores .
Slag cement and alumina-bearing pozzolans have an additional beneft as the hydration products are highly e ffective
in binding chloride ions, preventing further penetration into
the concrete (Thomas et al. 201 2). This effect is particularly
Getsuch
more
from
important in applications
as FREE
parkingstandards
garages, bridge
decks, and marine construction, where the reinforcing steel
is vulnerable to chloride-induced corrosion.
The indiscriminate use of SCMs is not necessarily benefcial.
Their
pozzolanic
and
hydraulic
reactions
take
time,
f
to
produce CH to participate in the reaction, and partly because
S CMs may vary signifcantly in terms o f kinetics or rate
of reactivity. The engineer must consider the properties of
the concrete at early ages. Because the pozzolanic reactions typically proceed more slowly than the hydration of
cement, extended moist curing is necessary to achieve the
best results.
Supplementary cementitious materials can often mitigate
the deleterious effects of elevated-temperature curing. Cao
and Detwiler (1 995) found that both silica fume and slag
were e ffective in refning the pore s tructure. Campbell and
Detwiler (1 993) tested a series of steam-cured concretes in
which the proportions of the various cementitious materials
were varied. They found that the total charge passed in 6
hours using AASHTO T277 varied by two orders of magnitude, with CSA Type 1 0 (now designated Type GU) portland
cement alone performing the worst, and optimized blends of
cement, slag, and silica fume performing the best. They did
partly
not us e
because
f
the
cement
mus t
hydrate
s u fciently
y as h in their s tudy.
Bentz et al. (1 995) showed that silica fume particles both
reduce the initial thickness of the interfacial transition zone
and react to convert CH to C-S-H. Fly ash has a similar,
but less-pronounced, effect due to its larger particle size and
lower pozzolanic activity.
3.4—Measurement o f transport properties
Measurement of the transport properties of concrete is
complicated by the interactions between the concrete and
the substance that is moving through it; the changing properties of concrete with time; and the sensitivity of transport
properties to variations in moisture, temperature, and other
conditions. Because many tests accelerate the transport
mechanis m o f the f uid or ion in question, they may induce
different or additional mechanisms of transport than what
would occur in service. They often make it impossible to
achieve the steady-state conditions that form the basis of
the various equations used to describe mass transport, and
may invalidate the as s umption o f laminar f ow used in many
calculations (Eq. (3.2.1 )). Furthermore, laboratory tests are
often conducted under highly controlled conditions that may
not accurately
re
f
ect actual
s ervice
conditions .
When
us ed
judiciously, however, the tests described in the following
may be helpful in comparing the suitability of different
concrete mixtures for a given exposure condition, or for
quality assurance/quality control purposes during construction (Puerto Rico DOT SP934). The following sections
discuss available ASTM tests used to characterize transport
properties, along with commonly used variations.
3.4.1 Ions
3.4.1.1 Coulomb test (ASTM C1202/AASHTO T277) —
Standard tests of ion transport focus on the penetration of
chloride ions into concrete. The most common test used for
this purpose is ASTM C1 202 (or AASHTO T277), in which
Standard
Sharing
Groupwith
anda our
chatsdiameter of 4 in.
a cylindrical
specimen
nominal
(1 00 mm) and a length of 2 in. (50 mm) is vacuum saturated with water before being placed in a test cell. The cell
contains a 3 percent solution of sodium chloride on one side
and a 0.3 N solution of sodium hydroxide on the other side.
An electrical potential of 60 V dc is applied for 6 hours. The
total charge passed during the test period is an indirect indication of the chloride ion penetrability of the concrete.
Essentially, ASTM C1 202 (or AASHTO T277) uses the
electrical conductivity of the concrete as a rapid index test
or surrogate for diffusivity. The main objections to the use of
this test method stem from the indirect nature of the measurement. While ion diffusion depends primarily on the microstructure and chemical binding capacity of the matrix, electrical conductivity depends on both microstructure and pore
solution chemistry (Buenfeld and Newman 1 987). Different
proportions of SCMs can profoundly affect the pore solution chemistry. For example, Page and Vennesland (1 983)
found that 1 0 percent silica fume, by substitution, reduced
the concentrations of Na+, K+, and Ca2+ by approximately
50 percent and that of OH – by approximately 75 percent in
the pore solution. Detwiler and Fapohunda (1 993) compared
the results of AASHTO T277 to those of a direct measure of
chloride ion migration for portland-cement concretes with
and without slag cement, and found that AASHTO T277
unduly favored the concretes containing SCMs. They attributed the differences between the two sets of test results to
differences in electrical conductivity due to differences in
pore solution chemistry.
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
Corrosion inhibitors such as calcium nitrite add ions to
the
pore
s olution
that
increas e
the
pore
f
uid
conductivity
and the charge passed. There is some evidence that calcium
nitrite increases the penetrability of the concrete matrix.
That is, increased coulomb values may be due to microstructural effects as well as changes in the pore solution
chemistry (Ann et al. 2006; Reou and Ann 2008 ). If ASTM
C1 202 (or AASHTO T277) is used in qualifying concrete
mixtures for use in construction, typically the concrete will
be tested without the corrosion inhibitor to qualify it, and
then with the corrosion inhibitor to provide a baseline value
for quality-control purposes. The latter measurement will be
considerably higher.
The relationship between electrical conductivity and
diffusion may also vary with the mechanism and type of
diffusion. Diffusion through voids and cracks differs from
bulk di ffus ion; thus , the pres ence o f f aws in the concrete
can s ignifcantly affect the res ults o f the tes t. In addition,
the measurement is taken before steady-state conditions are
reached (Zhang and Gjørv 1 991 ).
Low- quality
concrete
is
di ffcult
to
evaluate
properly
because the temperature of the specimens rises when current
is applied, increasing the rate of diffusion. McGrath and
Hooton (1 999) propos ed a modifcation to reduce the tes t
period to 30 minutes to eliminate this problem. High-quality
concretes may be di ffcult to dis tinguish from one another
because the total charge passed is so low, and because the
test results are variable (Hooton 1 989).
ASTM C1 202 (or AASHTO T277) has been criticized for
many of the previously-cited interferences (Andrade 1 993 ;
Feldman et al. 1 994; Streicher and Alexander 1 994; Shi 2004).
Despite its limitations, ASTM C1 202 (or AASHTO T277)
is rapid, convenient, and according to Hooton (1 989), whatever property it is measuring probably is coincident with
permeability. If properly interpreted, it can be used effectively for quality control in construction (Bognacki et al
201 0 ) , although it s hould not be used to make fne dis tinc tions among concretes and cannot be used to compare
concretes made with different materials or mixture proportions, or both (Shi 2004).
3.4.1.2 Ponding (ASTM C1543; AASHTO T259) and
bulk diffusion (ASTM C1556) —Another standard test that
is sometimes used is ASTM C1 543 (or AASHTO T259),
which involves ponding three concrete slabs at least 3 in. (75
mm) thick and a surface area of 46 in. 2 (0.030 m2) with a 3
percent sodium chloride solution for 90 days. The sides of the
slab are sealed and the bottom exposed to a drying environment at 50 percent relative humidity. If desired, the exposure
period can be extended to 6 months or 1 year. At the end of
the exposure period, the excess solution and salt buildup are
removed. Half-inch (1 2 mm) thick samples of the concrete
can be taken at two or three depths and analyzed for chloride
ion content, which is compared to a baseline value determined
on a companion concrete specimen not exposed to external
chlorides. Alternatively, the concrete slab can be sampled and
tested according to ASTM C1 556, in which a core from the
slab is milled or sliced to obtain samples at eight depths for the
purpose o f determining the apparent chloride di ffusion coe ff -
9
cient using Fick’s second law. This test can also be conducted
using a different salt in the ponding solution. Note that the
type of cation(s) present affects the rate of ingress of chloride ions because charge balance must be maintained, and the
associated cation(s) diffusing at a slower velocity will impede
the movement of chloride ions.
One of the most common objections to the use of ASTM
C1 543 is its duration. As the specimens must be cured for
1 4 days and then dried for 28 days before the beginning of
the ponding, the 90-day version of the test takes 11 8 days,
or longer in the case of extended curing, to conduct, after
which the samples must be analyzed. Most test programs for
high-performance concretes would use a ponding period of
at least 1 80 days.
Although the ponding test does provide a crude onedimens ional profle o f chloride ion ingres s , the profle is not
a re f ection o f chloride di ffus ion alone. The initial mode o f
ingress of the ions is by sorption into the dried concrete. The
exposure of the bottom face to a 50 percent relative humidity
environment during the test induces vapor transmission from
the wet front on the top surface to the dry bottom surface,
and chloride ions penetrate by wick action (Buenfeld et al.
1 995 ). Diffusion of the chlorides also takes place. McGrath
and Hooton (1 999) observed that, while all three of these
mechanisms do occur in bridge decks, the test exaggerates
the importance of the sorption component.
The 28-day drying period before the ponding begins
s ignifcantly
increas es
the
apparent
di ffus ion
coe ffcient,
especially for concretes containing SCMs (Ngala and
Page 1 997 ) . F or high- quality concretes , it may be di ffcult
to
develop
a
chloride
profle
bas ed
on
a
9 0 - day
ponding
period because so little chloride penetrates into the concrete.
Extending the ponding period to 1 80 days and increasing the
number of samples taken help to resolve this problem (Berke
and Hicks 1 992; Andrade and Whiting 1 996; Sherman et al.
1 996; McGrath and Hooton 1 999).
Precision o f the s ampling can make a s ignifcant di fference in the conclusions drawn from the results. In analyzing
their data from AASHTO T259 (a predecessor of ASTM
C1 543), McGrath and Hooton (1 999) showed that imprecise
sampling makes it di ffcult to dis tinguis h between a highquality concrete in which there is a high concentration of
chlorides near the surface, and a low-quality concrete in
which the chlorides penetrate much farther. Precision of the
sampling s pecifed in AS TM C1 5 5 6 avoids this problem.
ASTM C1 556 avoids some of the problems associated
with ASTM C1 543, as the concrete specimen is sealed on
all s ides except the fnis hed s ur face, and is s aturated with
a CH solution before exposure to the sodium chloride solution. Thus, the chlorides penetrate into the specimen only
by diffusion, not by sorption. The specimen is then placed
in a concentrated (1 65 g/L) sodium chloride solution for at
least 35 days. Longer exposure times are recommended for
mature concretes, concretes with low w/cm , or high-performance concretes containing SCMs. In practice, the exposure
time could be extended to 90 days or longer.
The sampling is more precise than for ASTM C1 543,
as indicated previously. A nominal 4 in. (1 00 mm) diam-
10
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
eter core from the slab is mounted in a mill or lathe and a
series of thin layers ground off. The dust from each layer is
collected separately and analyzed for acid-soluble chloride
content.
The apparent
di ffus ion coe ffcient is determined
using
nonlinear
regres s ion
analys is
to
ft the
data
to
F ick’ s
second law.
The use of Fick’s second law to reduce the data is a convenient but questionable practice. Pettersson (1 994) noted that
the applicability o f F ick’s s econd law, which is a simplif cation of a more general equation describing ion transport,
depends on the validity of three assumptions:
a) The material in which diffusion takes place is permeable and homogeneous.
b) The diffusion properties of the material do not change
with time or with the concentration of the diffusant.
c) No chemical reaction or physical binding of the diffusant occurs.
Pettersson (1 994) further noted that all three of these
assumptions are violated in the diffusion of chloride ions
through concrete. That is, concrete is heterogeneou s, its diffusion properties change with time and with the concentration
of the diffusant, and both chemical reactions and physical
binding can occur. ASTM C1 556 uses the term “apparent
chloride di ffus ion coe ffcient” to make clear that the result
obtained is not a true di ffus ion coe ffcient. Although AS TM
C1 556 in some cases may be conducted in less time than
ASTM C1 543 , it is still a lengthy test.
3.4.2 Fluids — One test for the absorption of water by hardGet
more
FREEa piece
standards
from
ened concrete is ASTM
C642,
in which
of concrete
at least 350 mL in volume is oven dried to constant mass and
then immersed in water until it again reaches constant mass.
The specimen is then boiled for 5 hours, allowed to cool, and
the mass determined again. The absorption after immersion
and the absorption after immersion and boiling are determined. This test is a measure of the absorption of the bulk
concrete. Note that oven drying may induce cracking in the
specimen, thus increasing the measured absorption.
ASTM C1 585 measures the water sorptivity, which is the
rate of absorption, of a concrete surface, which is often of
greater interest than the bulk concrete. A cylinder or core
4 in. (1 00 mm) in diameter and 2 in. (50 mm) in length
is conditioned to an internal relative humidity of 50 to 70
percent and then sealed on all but one surface. The mass of
the specimen is determined initially and after being placed in
contact with water. The mass is determined at close intervals
initially and at longer intervals up to an exposure time of 7
days, after which one additional measurement is taken. The
initial slope of the absorption-versus-time curve is taken as
the rate of absorption.
Abbas et al. (1 999) measured the permeability of concrete
to oxygen, which is easier than measuring its permeability to
water. Their calculations were based on Darcy’s Law, which
as s umes laminar f ow. This s impli fying ass umption is not
strictly
true
becaus e
in
very
s mall
pores
the
f
ow
is
partly
molecular in nature. Gas permeability varies with the degree
o f saturation o f the concrete; the coe ffcient o f permeability
varied over two orders of magnitude as a function of the
degree of saturation. Acceptable limits on the permeability
for durability vary depending on the exposure conditions
and performance requirements.
3.5 — Obtaining durable concrete
Obtaining durable concrete for given conditions of expofor the f uids
and ions of interest. Proper attention to all aspects of good
concrete practice is important. The mos t s ignifcant factors ,
however, are an appropriately low w/cm ; judicious use of
SCMs; and good workmanship, including mixing, placement, compaction, and curing. Elevated curing temperatures
can be deleterious to the transport properties of concrete,
although the use of an appropriate combination of cementitious materials can largely mitigate this effect. Optimization
of the concrete mixture proportions should be done using
the curing regime anticipated on the job and a test method
that bears some relation to the anticipated exposure conditions. In particular, compressive strength is not a surrogate
for durability.
Proper attention to control of cracks is also important;
there is little to be gained from concrete of low permeability
between the cracks. Good aggregate grading to minimize the
paste content, control of temperature and moisture conditions, and appropriate structural design and detailing can
minimize the width of cracks. Further guidance is available from the Transportation Research Board (2006) and
Detwiler and Taylor (2003).
sure requires s uitable mas s trans port properties
Standard
Sharing4—FREEZING
Group and our
chats
CHAPTER
AND
THAWING
CONCRETE
4.1 —Introduction
OF
Deterioration of concrete exposed to freezing can occur
there is s u ffcient internal mois ture that can freeze at
the given exposure conditions. The source of moisture can
be either internal or external. Internal is water that is already
in the pores of concrete that is redistributed by thermodynamic conditions to provide a su ffcient degree o f s aturation
at the point of freezing to cause damage. External is when
the water enters the concrete from an external source, such
as rainfall). Dry concrete (generally below approximately 75
to 80 percent internal relative humidity) is normally immune
to damage from freezing.
Young concrete can be damaged by a single freeze (4.2.2).
Mature concrete may be able to withstand repeated cycles of
freezing and thawing without damage. Thus, concrete that is
properly cured and reaches s u ffcient maturity be fore being
expos ed to freezing, such as concrete for columns or f oor
slabs, may tolerate freezing from exposure to a single winter
season before it is protected from the elements. Similar
concrete that is not properly cured and is exposed to freezing
conditions at an early age, such as sidewalks and exposed
foundation walls, may show deterioration after a few years
of exposure to repeated cycles of freezing and thawing.
The
s e ve rity
o f exp o s ure
s ho uld
be
quanti fe d
by
a
combination of freezing, which is the number of annual
cycles of freezing plus average low temperature reached
when
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
during each cycle, and moisture condition before each
cycle of freezing (4.2.6.4).
4.1.1 Concrete made with durable aggregate —Concrete
made with aggregate that is resistant to freezing and thawing
is primarily protected from damage through the use of
entrained air in the concrete mixture, with secondary protection from steps taken to limit the fractional volume of freezable water in the concrete. A description of damage due to
freezing and thawing of concrete made with durable aggregates, along with protection methods, is given in 4.2.
4.1.2 Concrete made with frost-susceptible aggregate —
A properly proportioned concrete mixture that has received
adequate curing can suffer damage from freezing and thawing
if it contains an aggregate (generally the coarse aggregate)
that is susceptible to damage from freezing and thawing.
A description of damage in concrete made with aggregates
susceptible to freezing and thawing, along with a description
o f aggregate identifcation procedures, is given in 4. 3 .
4.2—Frost attack o f concrete made with durable
aggregates
4.2.1 Description o f frost damage
4.2.1.1 Damage at early ages —Concrete in the early
stages of hydration ordinarily contains a considerable
amount of freezable water and has little or no tensile strength
to resist pressures due to freezing. Concrete that is allowed
to freeze under these conditions will develop ice lenses
approximately parallel to the surface exposed to freezing.
Additional ice lenses can develop under coarse aggregates.
When the concrete thaws and hydration resumes, the space
previously occupied by the ice lenses will form weak planes
that are susceptible to delamination or surface scaling.
4.2.1.2 Damage in cured concrete
4.2.1.2.1 Surface scaling —The most common form of
damage from freezing and thawing in hardened concrete is
surface scaling, that is, the loss of paste and mortar from the
surface of the concrete. Generally, layers less than 0.04 in.
(1 mm) thick are lost, but repeated cycles of freezing and
thawing can lead to removal of additional material. Scaling
is cons iderably accelerated by deicing s alts. Vehicle tra ffc
or other surface contact can also accelerate scaling by
aiding in the removal of loosened material. Consequences
of scaling include change in appearance; change in surface
smoothness; and, in severe cases, loss of concrete cover over
reinforcing steel.
4.2.1.2.2 Internal deterioration —Though less common,
internal deterioration can have more severe consequences
than surface scaling due to the loss of structural integrity
of the concrete. Internal deterioration manifests itself as a
loss of strength in the paste of the concrete. Modern concrete
practice has practically eliminated this form of damage from
freezing and thawing by requiring a proper air-void system
and adequate curing be fore the frs t expos ure to freezing
temperatures.
4.2.2 Preventing frost damage in new concrete
4.2.2.1 Protection from early freezing —Young concrete
should be protected from freezing by following the procedures and maintaining the minimum temperatures recom-
11
mended in ACI 306R. A fter cons olidation and fnis hing, the
concrete should be protected from cooling too rapidly by the
use of insulated forms, curing blankets, and other procedures
described in ACI 306R. Allowing the concrete to cool too
rapidly could result not only in early freezing, but also in
thermal cracking of the concrete (ACI 306R; ACI 308R).
4.2.2.2 Minimum curing be fore freezing —Adequate
curing, including preventing excessive drying and maintaining adequate temperature, will ensure that the concrete
has hydrated s u ffciently to s ubs tantially reduce the amount
of freezable water. A recommended minimum strength
that should be attained before the concrete temperature is
allowed to drop below freezing is 500 psi (3.5 MPa) (Powers
1 962a). Once this strength has been achieved, a single freeze
will generally not permanently damage the concrete (ACI
308R). If repeated cycles of freezing and thawing are anticipated, the concrete should be kept warm long enough to
allow it to develop a compressive strength of at least 3500
psi (25 MPa) if it will not be exposed to deicing salts, and
4500 psi (32 MPa) if it will be exposed to deicing salts. The
strengths of 3500 and 4500 psi are average in-place strengths
needed before the concrete is exposed to repeated cycles of
freezing and thawing. Table 1 9.3.2.1 in ACI 31 8-1 4 refers
to
the
s pecifed
design
strength.
The
two
strengths
are
not
the same, as concrete could be exposed to repeated cycles of
freezing and thawing at an age s ignifcantly earlier than the
age as s ociated with the s pecifed des ign s trength.
4.2.3 Preventing frost damage by proper design —Much
concrete now in service has withstood repeated cycles of
freezing and thawing for many years. While some of this
concrete has remained undamaged because it was never
allowed to contain enough freezable water to cause damage,
most of it has remained durable because proper precautions
were taken to avoid such damage (Mather 1 990). The three
most important precautions to provide resistance to freezing
and thawing are discussed in 4.2.3.1 through 4.2.3.3.
4.2.3.1 Reducing freezable water—The likelihood of
damage from freezing and thawing is reduced by decreasing
the amount of freezable water in concrete. For conventional
mixtures, this has generally been accomplished by lowering
the w/cm , which is a maximum of 0.50 for moderate exposure and 0.45 for severe and very severe exposure (Table
4.2.3.1 a), combined with adequate curing to ensure a
minimum compressive strength of approximately 3500 psi
(25 MPa) before exposure to repeated cycles of freezing and
thawing (4500 psi [32 MPa] if deicing salts are present).
Note that for instances in which corrosion is a concern, a
lower w/cm may be required if deicing salts are present (ACI
31 8-1 4).
Limiting the w/cm to a s pecifed maximum has the e ffect
of reducing the amount of freezable water in the cured
concrete initially. Requiring a minimum strength before
freezing helps ensure that the tensile strength of the paste is
su ffcient and the fractional volume that could be occupied
by freezable water in saturated concrete has been adequately
reduced by the formation of hydration products.
Modern concrete mixtures may contain admixtures, additives, and supplementary cementitious materials (SCMs)
12
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
Table 4.2.3.1 a—Freezing-and-thawing exposure
classes
Table 4.2.3.1 c—Cementitious materials limitations
for Exposure Class F3b
Exposure
Maximum percent of total
Class
Severity
Condition
Cementitious materials
cementitious materials by mass *
F0
Not
applicable
Concrete not exposed to freezing conditions
Fly ash or other pozzolans
conforming to ASTM C61 8
25
Slag conforming to ASTM C989/
C989M
50
Moderate
Concrete exposed to freezing and thawing
conditions, but very low probability of
concrete being near saturation at time of
exposure *
Silica fume conforming to ASTM
C1 240
10
Total o f f y ash or other pozzolans,
slag, and silica fume
50 †
Total o f f y ash or other pozzolans
and silica fume
35 †
F1
F2
Severe
F3
Very severe
Concrete exposed to freezing and thawing
conditions, with a high probability of
concrete being near saturation at time of
exposure, but no deicing chemical exposure †
Concrete exposed to freezing and thawing
conditions as well as deicing chemicals ‡
*Examples are vertical surfaces above the level of snow accumulation or horizontal
elevated f oors in areas protected from direct exposure to moisture.
†
Examples are: vertical surfaces below the level of snow accumulation; vertical
surfaces with suffcient moisture exposure to allow the concrete to be near saturation
prior to freezing; retaining walls or other vertical elements with one side exposed to
moisture; and slab-on-ground that is not protected from freezing.
‡
Examples are: vertical surfaces that may have deicing-chemical-contaminated snow
piled against them; sidewalks or pavements that receive deicing chemicals; and
concrete that received frequent exposure to seawater as well as freezing-and-thawing
conditions.
Table 4.2.3.1 b—Requirements by freezing-andthawing exposure class
Minimum
Limits on
Exposure
fc , * MPa
Class
(psi)
F0
No
restriction
No
restriction
No
restriction
No restriction
F1
25 (3500)
0.50
Table
4.2.3.2.4
No restriction ‡
F2
25 (3500)
0.45
Table
4.2.3.2.4
No restriction ‡
F3a§
32 (4500)
0.45**
Table
4.2.3.2.4
Table 4.2.3.1 c ‡
F3b #
32 (4500)
0.45**
Table
4.2.3.2.4
No restriction ‡
Maximum
cementitious
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from
w/cm
Air content
materials
†
*
The minimum average compressive strength that should be achieved before initial
exposure to freezing and thawing.
†
The maximum w/cm for the in-place concrete to provide adequate restriction of freezable water in the properly-cured concrete.
‡
High cementitious material replacement for portland cement frequently results in
lower rates of strength gain. Care should be taken to ensure that adequate curing (moisture, temperature, and time) is provided so that the minimum fc is achieved before
initial exposure to freezing and thawing.
§
Hand-fnished surfaces.
#
Formed and machine-fnished surfaces.
**
A lower
/
w cm
may be needed when corrosion is of concern (ACI 31 8-1 4).
that permit reduction of the amount of freezable water in
concrete. Slag cement or pozzolans can refne the pore struc ture at a given w/cm , resulting in a smaller fraction of the
porosity containing freezable water. The use of slag cement
or pozzolans has the added benefcial e ffect o f reducing
the rate that water can penetrate through the concrete. This
means that water removed by evaporation or hydration will
be replaced more slowly when the concrete is exposed to
*
The total cementitious materials also include ASTM C1 50/C1 50M, ASTM C595/
C595M, ASTM C845/C485M, and ASTM C11 57/C11 57M cements.
The maximum percentage should include:
(a) Fly ash or other pozzolans in Type IP blended cement, ASTM C595/C595M or
ASTM C11 57/C11 57M
(b) Slag used in the manufacture of an IS blended cement, ASTM C595/C595M or
ASTM C11 57/C11 57M
(c) Silica fume, ASTM C1 240, present in a blended cement
†
Fly ash or other pozzolans and silica fume shall constitute no more than 25 and 1 0
percent, respectively, of total mass of the cementitious materials.
water. Numerous feld observations have indicated when
surfaces are fnished by hand, as opposed to machinefnished or cast against formwork, and exposed to deicing
salts (Exposure Class F3a in Table 4.2.3.1 b), limits on the
SCMs may
be required
dueand
to a variety
of factors that appear
Standard
Sharing
Group
our chats
to include modifcation o f the entrained air-void system
(4.2.3.2) as well as superfcial changes to the w/cm . Table
4.2.3.1 c gives recommended cementitious materials limitations for Exposure Class F3a (Thomas 1 997).
Some researchers have hypothesized that it may be
possible to produce concrete with so little freezable water
that the concrete would not be damaged on freezing. Hooton
(1 993) and Pigeon (1 994) found that the cost and di ffculty
in placing and fnishing such low w/cm concrete would make
it impractical.
4.2.3.2 Entrained air-void system —Resistance to freezing
and thawing of a concrete mixture is substantially improved
by incorporating entrained air voids into the concrete. To
achieve maximum effectiveness, these air voids should
be evenly distributed throughout the paste portion of the
concrete. Their spacing should be close enough to prevent
the development o f suffcient pressures from freezing to
fracture the concrete.
Because voids reduce the strength and stiffness of most
concrete mixtures, there is a natural tendency to limit
entrained air in concrete. An adequate air-void system,
however, is necessary for resistance to freezing and thawing.
The specifc parameters normally used to evaluate an airvoid system along with the generally accepted minimum
(or maximum) values are described (4.2.3.2.1 through
4.2.3.2.4).
4.2.3.2.1 Spacing factor L —Spacing factor L is an
approximation of the average distance from anywhere in the
cement paste to an air void. The following assumptions are
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
made for this parameter: the voids are spherical, of equal size,
and evenly distributed in a simple cubic lattice throughout
the paste (Powers 1 949). The method for determining this
parameter is described in ASTM C457/C457M . The range of
spacing factors is generally from 0.004 in. (0.1 mm) or less to
approaching 0.04 in. (1 mm) for mixtures that do not contain
entrained air. The generally accepted maximum spacing
factor value for concrete with good resistance to freezing and
thawing is approximately 0.008 in. (0.20 mm) (Powers 1 954;
Backstrom et al. 1 954, 1 958b). Other researchers (Dubovoy
et al. 2002; Attiogbe et al. 1 992) have shown that concrete
with spacing factors greater than 0.008 in. (0.20 mm) may
exhibit satisfactory freezing-and-thawing resistance under
specifc conditions in laboratory tests. CSA A23 .1 -1 4/CSA
A23.2 requires an average spacing factor of less than 0.009 in.
(0.23 mm) with no single value to exceed 0.01 0 in. (0.26 mm).
This is relaxed to an average value of 0.0098 in. (0.25 mm) for
concrete with specifed strength greater than 1 0,000 psi
(70 MPa).
4.2.3.2.2 Specifc surface α —Specifc surface α is a
measure of the surface area per unit volume of voids. The
method for determining this parameter is described in ASTM
C457/C457M. A basic assumption o f specifc surface is that
all voids are spherical, which makes it a function of average
chord length alone (Powers 1 949). Specifc surface is, there fore, a good indicator of average void size. As average void
size goes up, specifc surface goes down. Smaller voids
provide compliance with the L requirement at lower air
contents. The generally accepted value o f specifc surface
for resistance to freezing and thawing is a minimum of
600 in. 2 /in. 3 (25 mm 2 /mm 3 ) (Powers 1 949), though other
researchers have shown that concrete with a specifc surface
area less than 600 in. 2 /in. 3 (25 mm 2 /mm 3 ) may exhibit satisfactory freezing-and-thawing resistance in laboratory tests
(Dubovoy et al. 2002; Attiogbe et al. 1 992).
4.2.3.2.3 Philleo factor F′—Philleo (1 955) developed an
air-void parameter as a means to eliminate the assumptions
made for the spacing factor—namely, that all voids are of
equal size and spacing. His equation is based on the assumption that voids are randomly sized and distributed and establishes a relationship between the air-void distribution and
the percentage of paste that is within a given distance of an
air void. Philleo used the work of Willis and Lord (1 951 ) to
establish a relationship between air-void chord lengths and
voids per unit volume. The equation is used to either determine the percentage of paste that is protected because it is
within a specifed distance o f an air-void, or alternately, to
determine the distance from an air-void resulting in a specifed percentage o f the paste. This distance, called the Philleo
factor, is often compared with the spacing factor. In actuality, however, the Philleo factor is more sensitive to the true
air-void distribution than the air content, paste content, and
number o f voids that inf uence the spacing factor.
This parameter has not been widely accepted as a measure
of resistance of concrete to freezing and thawing, primarily
due to the di ffculty in acquiring the data necessary for
calculation. These data consist of a record of all chord
13
lengths measured in a linear traverse (ASTM C457/C457M).
Most petrographers, however, use the alternate modifed
point count method (ASTM C457/C457M), which does not
collect the necessary data. While no specifc criterion for the
maximum F′, which is the distance for a given percentage
of paste protected, has been determined, an examination
of a considerable amount of linear traverse (ASTM C457/
C457M) data for a number of concrete specimens having
spacing factors of approximately 0.008 in. (0.20 mm) and
specifc surface values o f approximately 600 in. 2 /in. 3 (25
mm 2 /mm 3 ) suggests that a maximum acceptable distance
between an air void and 90 percent of the paste, P90 ′, should
be approximately 0.002 in. (0.04 mm) (Janssen and Snyder
1 993 , 1 994).
4.2.3.2.4 Air content—The aforementioned air-void
parameters, while excellent indicators of the protection from
freezing and thawing provided by the air-void system, are
di ffcult to measure in the feld. Total air content is there fore
generally specifed and measured. Total air content includes
both the entrained air voids and the larger air voids that are
not removed by consolidation.
The use of an air-entraining admixture complying with
ASTM C260/C260M can provide a proper system of
entrained air-voids when a specifed total air content o f
the concrete mixture is achieved. The actual air content
necessary to ensure the production of the necessary airvoid system is affected by mixing action, workability of
the mixture, cement composition, types and amounts of
other admixtures, and others (Whiting and Stark 1 983 ). In
addition, concrete handling during transport, placing, and
fnishing can affect the entrained air-void system (4.2.4).
Recommended air contents for fresh concrete are given
in Table 4.2.3.2.4. These recommendations consider the
higher air requirements of concrete mixtures with higher
paste contents, which would be smaller nominal maximum
aggregate sizes, as determined by Klieger (1 952, 1 956) and
the severity of exposure; higher exposure severity increases
the probability of damage from freezing and thawing and,
therefore, demands greater protection. The values shown are
general recommendations; local conditions and experience
with specifc mixtures, admixtures, and construction proce dures could warrant other values.
Achieving the total air content specifed in Table 4.2.3.2.4
does not always ensure frost protection of the paste. Rather,
a mixture should achieve the minimum and maximum
values for the air-void parameters discussed previously. In
most cases, the minimum air contents from Table 4.2.3.2.4
will achieve the necessary air-void parameters.
4.2.3.3 Design details —Physical details that allow for the
repeated wetting or restricted drying of concrete surfaces
should be avoided. Examples of these include roof driplines on sidewalks, and structural members directly below
unsealed joints of bridge slabs. Be careful to always provide
positive drainage o f runo ff from f at areas by methods
such as sloping surfaces, which is typically 2 percent.
Special precautions should be taken when runoff is likely to
contain deicing salts or other aggressive chemicals to avoid
increasing the risk of damage to concrete surfaces.
14
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
Table 4.2.3.2.4—Recommended air contents
Air content, percent*
Nominal maximum
aggregate size, in.
Exposure Class F2
(mm)
Exposure Class F1
and F3
3/8 (9.5)
7
7.5
1 /2 (1 2.5)
7
7
3/4 (1 9)
6.5
7
1 (25)
6.5
6.5
1 -1 /2 (37.5)
6
6.5
2 (50)
6
6
3 (75)
5
5.5
*
Field tolerance on air content is recommended as ±1 -1 /2 percent. Air content recommendations are based on 1 8 percent air in the paste portion of the concrete with a
Vinsol resin air-entraining agent (from an analysis of work by Klieger [1 952]). Mixture
proportions based on guidance in ACI 211 .1 for angular coarse aggregates along with
the maximum w/cm values from Table 4.2.3.1 b were used to determine the air content
recommendations. Mixtures using rounded aggregates will require approximately 1
percent less air due to the lower paste contents associated with rounded aggregates.
Consideration should also be given to structures that allow
water contact with the side away from freezing, such as water
tanks. In many cases, the evaporation rate on the outside of
the tank will exceed the rate that water can migrate through
the tank wall to replenish the evaporation. In these instances,
the amount of freezable water in the portions of the concrete
that freeze will be below critical degree of saturation. In
some cases, however, water migration to the freezing surface
can exceed the evaporation rate, leading to potential damage
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from freezing and thawing.
In very
cold climates,
freezing
can extend through the concrete wall, resulting in severe
damage. Low rates of moisture movement and consideration
of the depth of freezing are necessary in the design of these
types of structures (ACI 350).
4.2.4 Preventing frost damage by proper practice —A properly proportioned concrete mixture can still suffer damage
from freezing and thawing if either the air-void system or
the amounts of freezable water are adversely affected during
construction. The effects of construction practices on maintaining the quality of the concrete mixture are described in
4.2.4.1 through 4.2.4.4.
4.2.4.1 Transporting and placing —The air voids that
provide protection against freezing-and-thawing damage
are produced during mixing (Whiting and Stark 1 983 ). Airentraining admixtures stabilize the bubbles that are produced
during mixing, but do not generate them. Therefore, mixing
is critical in forming and distributing the bubbles throughout
the mixture. Upon delivery and discharge of the concrete, the
number, volume, and size distribution of the bubbles depend
on the mixing action and the degree to which the particular
combination of ingredients has worked to generate, stabilize, and retain the bubbles.
The system of air bubbles in the fresh concrete at placement will be sensitive to many factors, including the type
and effectiveness of the mixer, duration of mixing, mixing
speed (rpm), haul time, batch size, and general condition of
the mixing equipment. These factors affect the time required
for the air-entraining admixture to stabilize the bubbles.
Fast-acting air-entraining admixtures may perform well
even with short mixing periods or brief haul times, whereas
more slowly-acting admixtures can perform better with a
longer mixing or haul time. The degree that air is incorporated in the fresh concrete during mixing depends on the
shearing, tumbling, or wave-breaking action of the concrete
in the mixer itself. This action will depend on the slump
o f concrete
and cleanliness
and e ffciency o f the mixing
blades. Air incorporated in concrete is subdivided into
smaller bubbles by continued mixing and stabilized by the
air-entraining admixture, thus minimizing air loss. Immediately before being discharged, the freshly mixed concrete
contains air bubbles of a wide distribution of sizes and with
varying effectiveness in terms of providing frost resistance,
which all have their origin in the engulfment or entrapment
of air during mixing (Mielenz et al. 1 958a,b; Backstrom et
al. 1 958a,b).
Concrete is often placed by pumping from the delivery
truck to the formwork. Advances in pumping technology
have resulted in increased distances, lift heights, and delivery
rates. These advances have been accompanied by increases
in the pressure capacity of the pumps, with pressures in the
range of 300 to 500 psi (2 to 3.5 MPa) being fairly common
(Cooke 1 990). Unfortunately, concerns have developed
when the air content measured after pumping was not the
same as the air content before pumping. In most cases,
the air content had decreased after pumping, but in some
instances, it increased (Cooke 1 990; Dyer 1 991 ; Whiting
and Nagi 1 998 ). The possible loss of entrained air can be a
Standard
Sharing
our chats
major concern
for Group
concreteand
exposed
to severe freezing-andthawing environments.
Pumping concrete includes a number of activities that
could contribute to a change in the air content of the concrete.
During pumping, the concrete falls through a grating in the
pump hopper, is forced under pressure through a relatively
small-diameter pipe, moves through a series of bends, experiences changes in both elevation and pipe material, which is
generally steel to rubber, is then released from pressure as it
exits the pipe, and often falls some distance into the formwork. Pumping may also be continuous or interrupted at
various times. This sequence o f events makes it di ffcult to
isolate the mechanism(s) that induce(s) the loss or gain of air.
A number of mechanisms have been proposed to explain
the change in air content that sometimes occurs when
concrete is pumped. During pumping, a vacuum could form
in the line, especially at low pumping rates with a long
portion of the pumping being downhill, which would enlarge
and remove air voids. Air voids could also be lost when the
concrete falls through the grating in the pump hopper or
when it falls into the formwork after exiting the pump line
(Yingling et al. 1 992; Janssen et al. 1 995 ). While changes
in the air-void system have been documented in multiple
cases (Whiting and Nagi 1 998; Janssen et al. 1 995; Pleau
et al. 1 995 ), evidence of reduced resistance to freezing and
thawing has not been identifed ( Elkey et al. 1 993 ; Whiting
and Nagi 1 998).
4.2.4.2 Consolidating —On discharge of fresh concrete,
air
voids
could
be
entrapped
in
the
concrete.
While
flling
forms, it is virtually impossible to avoid the inclusion of air
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
pockets. When such large air pockets form during mixing,
continued mixing breaks them down to smaller sizes and
distributes them into the mixture. No such opportunity exists
once the concrete is cast. For this reason, it is necessary to
reduce the number and size of these pockets trapped in the
mixture by vibration.
Vibration accomplishes two purposes with regard to the
removal of air pockets in fresh concrete. First, vibration
can liquefy the concrete in the same manner as earthquakes
liquefy certain types of soils. Many of the large, buoyant air
pockets
that were
trapped
in the
f
f
f
s emi
to the s ur ace through the temporarily
uid mixture
can ris e
uid mixture. S econd,
vibration imposes a cyclic compression in the concrete
that locally increases, then decreases, the water pressure
surrounding the air bubbles. This causes the bubbles themselves to compress and decompress at the frequency of
the vibrator (Young 1 989). Air bubbles break when forced
to compress and decompress at a critical frequency that
varies with their size—the larger the bubble, the lower the
frequency. Conversely, the higher the vibration frequency,
the smaller the bubble that can be broken. Simon et al.
(1 992) showed that vibration at approximately 6000 vibrations per minute begins to break bubbles in the size range
of the so-called entrained air. Evidence also indicates that
some portion of the air from broken bubbles escapes, while
the rest is incorporated into other air bubbles in the mixture.
It takes time for the energy imparted by the vibrator to
liquefy the surrounding concrete and for the water pressure
to build up so as to compress the air bubbles. It also takes
time for the bubbles to rise to the surface, break, or both. For
these reasons, short intermittent vibration may have little
effect on the concrete in general or on the air bubble system.
Refer to ACI 309R for guidelines for appropriate vibration
procedures.
In s ummary, vibration o f fres h concrete refnes the airvoid system by encouraging the loss of larger air voids while
retaining the smaller ones. The exact distinction between
the size of bubbles removed and those left in place depends
on the concrete mixture and on the frequency, duration, and
intensity of vibration.
4.2.4.3 Finishing —After concrete is placed and consolidated, it is still possible to modify the air bubbles at the
surface of the fresh concrete, and therefore the air voids in
the hardened concrete, during fnis hing. Repeated pas ses o f a
fnis hing tool can force bubbles together, res ulting in fewer,
larger bubbles. The air content of the concrete at the surface
can be reduced by over- fnis hing,
that is, over- manipulation
of the surface can reduce resistance to freezing and thawing.
This is particularly the case i f the s ur face is fnis hed while it
is still covered with bleed water or if water has been applied
to make it eas ier to fnis h. F inis hing with free water on the
surface not only weakens that surface by increasing the w/
cm locally, but also increases porosity. When coupled with a
localized reduction in air content, resistance to freezing and
thawing is likely to be signifcantly reduced. Thomas (1 997)
found that this was especially true for hand- fnis hed concrete
having a high dos age o f S CMs (5 0 percent C las s F f y as h) .
15
and fnis hed areas o f the s ame concrete did
not show scaling damage.
For concrete slabs, it has been found that hard-troweling of
air-entrained concrete has the potential to blister or delaminate at the surface (ACI 301 ; ACI 302.1 R; Tarr and Farny
2008 ). Hence, as per ACI 301 , concrete for slabs to receive a
Machine- placed
hard trowel fnis h s hould not contain an air- entraining
agent
or have a total air content greater than 3 percent.
4.2.4.4 Curing —Curing is defned as the maintenance o f
a satisfactory moisture content and temperature in concrete
during its early stages so that the desired properties may
develop (ACI 308R). Overall, resistance to freezing and
thawing increases as continued curing develops microstructure and reduces the porosity and hydraulic conductivity
of the concrete. The result is a concrete that is less likely
to become critically saturated. Curing further increases the
compressive and tensile strength of the concrete, which
increases the resistance to pressure from freezing. These
attributes combine to produce a concrete that is less susceptible to freezing-and-thawing damage. The issue is made
more complicated, however, when concrete is cast in weather
where there is risk of freezing. In this situation, attempts
to cure the concrete to improve its overall durability can
con f ict with the fact that curing can increas e the ris k o f s atu rating concrete during exposure to early freezing. It is generally accepted that properly air-entrained concrete can sustain
one freeze cycle when it has attained a compressive strength
of 500 psi (3.5 MPa), and repeated freezing-and-thawing
cycles when saturated at a compressive strength (in-place) of
3500 psi (25 MPa). This means that curing procedures need
to be carried out to maintain moisture content to improve the
quality of the concrete, as well as measures to prevent early
freezing (ACI 306R; ACI 308R).
4.2.5 Preventing frost damage in existing concrete that
lacks adequate air-void system —Concrete that lacks an
adequate air-void system to protect it from the anticipated
exposure conditions is sometimes encountered. Protecting
the concrete from damage caused by freezing and thawing
then requires keeping it dry. Powers and Brownyard (1 947)
presented the thermodynamic calculations to show that
concrete dried to 85 percent relative humidity at room
temperature would contain no freezable water at –0.4°F
(–1 8°C). Therefore, concrete dried to an internal relative
humidity below approximately 75 to 80 percent would
rarely, if ever, contain freezable water.
4.2.5.1 Sealers —Concrete that can be adequately dried
can sometimes be kept dry by sealing the surface of the
concrete with some sort of protective barrier system to
prevent the reintroduction of moisture. Slab-on-ground and
similar construction would also require a vapor retarder
beneath the concrete to reduce the movement of water in
either liquid or vapor form from entering the concrete from
underneath. Details of various protective barrier systems can
be found in ACI 51 5.2R. Vapor retarders for use under slabs
are discussed in ACI 302.1 R.
In many cases, not all sides of a concrete member are
accessible for sealer treatments. Care should be taken so
that the sealed surface does not prevent the evaporation of
16
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
moisture that may have entered from unsealed surfaces. An
example would be the sealing of the top surface of a slabon-ground. The sealer could restrict evaporation of moisture
entering the slab from the bottom. The concrete could end
up wetter than if it had never been sealed. Vapor-permeable
sealers, which limit the intrusion of liquid water and permit
evaporation of moisture from the sealed surface, should be
evaluated for application-specifc conditions.
4.2.5.2 Drainage and other methods —The moisture in
concrete can o ften be suffciently limited to reduce the possi bility of damage due to freezing and thawing if attention is
given to the removal of water from the area of the concrete.
While methods of moisture reduction are generally application-specifc, the examples in 4.2.5.2.1 through 4.2.5.2.3 are
presented.
4.2.5.2.1 Drainage —Water allowed to pond on a concrete
surface can contribute signifcantly to the moisture in the
concrete. By providing adequate drainage to prevent ponding,
the water absorbed in the concrete can be minimized.
4.2.5.2.2 Maintenance —Designs that provide for rapid
drainage from concrete surfaces can be defeated by poor
maintenance. An example is the ponding of water on a
concrete slab being caused by leaves blocking a drain. Proper
design should be supported by adequate maintenance. Other
times, improper maintenance or other activities could result
in the unnecessary accumulation of water. Snow pushed off
a sidewalk and against a concrete wall could result in the
accumulation of moisture in the wall. As the snow against
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the wall melts, the remaining
snow
couldstandards
serve as a dam,
holding the water against the wall. Normally, a surface with
little moisture absorption, this vertical wall section could
absorb enough melted snow to lead to damage from subsequent freezing.
4.2.5.2.3 Redirection o f water fow—Downspouts
emptying water across a concrete slab can be redirected
so that runo ff does not f ow across the concrete. This can
reduce the moisture exposure of the surface enough to
reduce damage from freezing and thawing.
4.2.6 Theories for frost damage —Damage in concrete
from freezing and thawing occurs as internal damage or
surface deterioration (Cordon 1 966). Historically, the capillary-void system of the cement paste has been the focus of
most investigations. There is no consensus regarding the
mechanisms responsible for damage in cement paste. The
damage has been attributed to hydraulic pressure buildup
as water is forced away from the freezing front, to osmotic
pressure gradients driving water toward the freezing centers,
to vapor pressure potentials, and to combinations of these
processes (Powers 1 945 , 1 954, 1 955 , 1 975 ; Powers and
Helmuth 1 956; Helmuth 1 960a; Litvan 1 972; Penttala 1 998 ;
Setzer 1 999, 2002; Scherer and Velenza 2005 ; Coussy and
Monteiro 2008 , 2009). These are described in more detail in
the following sections.
4.2.6.1 Moisture expulsion —Powers proposed that ice
nucleated and grew in capillary pores, forcing them to dilate
or expel excess water from freezing sites. Elevated hydraulic
stresses would arise as water was expelled due to the relatively low permeability of cement paste. Distance from
the void boundary, the degree of saturation, and the rate of
freezing would inf uence the magnitude o f hydraulic-stress
buildup (Cordon 1 966; Powers 1 945, 1 954, 1 955).
4.2.6.2
Osmotic pressure —The hydraulic-pressure
concept was later modifed when experiments showed
signifcant evidence that moisture was moving toward, rather
than away from, freezing sites (Powers and Helmuth 1 956;
Helmuth 1 960a; Powers 1 975). Marchand et al. (1 995) and
Penttala (1 998) proposed that not all of the water in the capillary pores is freezable due its surface tension and the small
diameter of the pores containing it. Water in the largest voids
would freeze before the water in the smaller voids. When
the water in the larger voids freezes, the concentration of
the dissolved salts increases locally, causing a concentration
gradient in the pore solution. Water is thought to move from
the smaller voids to the larger ones to reduce this gradient.
The resulting f ow is thought to cause damage. Litvan (1 972)
proposed a similar theory, also founded on thermodynamic
arguments, but cast in terms of vapor-pressure gradients
between supercooled water and ice instead of salt concentration gradients.
Penttala (1 998), Scherer and Valenza (2005), Setzer
(1 999, 2002), and Coussy and Monteiro (2008, 2009) have
combined the moisture expulsion and osmotic pressure theories to account for the rate of freezing, degree of saturation,
dispersion of air voids, and the paste microstructure. These
theories agree that forces are exerted on the paste from
the movement of water from the small to the large voids.
Standard
Group
and
Scherer Sharing
and Valenza
(2005)
addour
thatchats
when the larger voids
fll with ice, local pressure from ice crystallization would be
expected. These pressures will increase with the shape and
curvature of the pores. A combination of these factors lead
to initial damage.
4.2.6.3 Ice lens growth —Whereas the previous mechanisms
may describe the deterioration in small saturated samples
frozen rapidly and cyclically in the laboratory, progressive
ice accretion in cracks during periods of sustained temperatures slightly below 32°F (0°C) may generally dominate the
further degradation of previously damaged or microcracked
concrete (Litvan 1 978 ). A theoretical model was developed
and proven successful for predicting freezing-and-thawing
damage in rock (Walder and Hallet 1 985 ; Hallet et al. 1 991 ).
In this model, freezing attack in homogeneous porous solids
is viewed as occurring in an open system where microcracks
are internally pressurized by ice accretion fueled by migration moisture from either liquid or vapor; the latter is in turn
induced by thermally driven free-energy gradients. This ice
accretion model, also known as the segregation ice model, is
consistent with Gilpin’s (1 980) study of freezing effects in
porous media as applied to soils.
4.2.6.4 Implications o f freezing-and-thawing damage
mechanisms —Uncertainty and opposing views of fundamental processes governing freezing in concrete undermine efforts to develop tests of the resistance to freezing
and thawing of concrete. Several testing strategies have
been used that involved relatively rapid cycling of freezing
and thawing in saturated and dry environments. In North
America, a freezing-and-thawing cycling test, ASTM C666/
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
C666M, is used to determine the resistance of concrete
mixtures to internal damage. ASTM C672/C672M evaluates
scaling resis tance.
The latter relies on vis ual class ifcations ,
sometimes supplemented with mass loss measurements.
These tests, like some of their predecessors, are criticized
for not adequately representing typical environmental conditions. The tests cycle between extreme temperatures too
rapidly. The CDF test (RILEM TC 1 1 7-FDC 1 996) provides
greater reproducibility and is more quantitative with respect
to scaling measurements. Hallet et al. (1 991 ) proposed that
conventional tests give only limited guidance for understanding
the
process es
governing
freezing at feld s ites
where the thermal or hydraulic regimes are very different
from those in the laboratory. Extrapolation from diurnal or
more
frequent freezing- and- thawing experiments to feld
conditions should be viewed with particular caution because
distinct physical processes may govern each. This situation
calls for consideration of diagnostic freezing-and-thawing
tests in which both processes can be distinguished and
probed systematically. This also suggests that considerably
more caution is needed when attempting to relate standard
laboratory test results to spacing factors and to other design
criteria for effective resistance to freezing and thawing in
service. Some concrete mixtures pass accepted laboratory
tes ts and do not per form well in the feld, while others fail
the tests and per form quite s atis factorily in the feld. This is
probably because the tests for internal damage and scaling
due to freezing and thawing do not addres s all o f the s ignif cant variables. Stark (1 989a) indicated that the potential
role of several key factors in freezing-and-thawing damage
has not been appreciated; these include the magnitude and
duration of exposure to sustained temperature and moisture
gradients , and the cumulative time o f exposure to specifed
temperature ranges.
4.3—Frost attack o f concrete made with
nondurable aggregates
Deterioration due to freezing and thawing of properly
proportioned, air-entrained concrete made with aggregate susceptible to freezing-and-thawing damage is often
referred to as D-cracking. Many types of coarse aggregate
have
been
identifed
as
s us ceptible
to
D - cracking,
while
other sources of the same kind of rock have not been found
susceptible. The pore structure of the coarse aggregate is
thought to be the primary contributing factor to susceptibility to D-cracking. Descriptions of the appearance and
development of D-cracking are presented in 4.3.1 , while the
prevention of D-cracking in new construction is covered in
4.3.2. Mitigation of D-cracking in existing construction is
given in 4.3.3. Theories and mechanisms of D-cracking are
covered in detail in 4.3.4.
4.3.1 Description o f D-cracking
4.3.1.1 General description —D-cracking is characterized by cracks through the coarse aggregate and mortar of
the concrete. Away from the cracks, both mortar and coarse
aggregate are strong and show no signs of deterioration.
The development of D-cracking requires considerable
moisture and repeated cycles of freezing and thawing. As
17
a result, D-cracking usually appears close to joints, cracks,
edges, and corners where moisture enters from more than one
surface. D-cracking generally appears as a series of cracks
approximately parallel to the primary moisture source.
4.3.1.2 Flatwork—The most common appearance of
D - cracking is in at- grade f atwork s uch as highway pave ments, parking lots, and sidewalks. These are areas that
frequently have readily available moisture from precipitation runoff, from multiple directions at joints and cracks,
and on the bottom of the slab unless there is an effective
vapor retarder beneath it. Often, the earliest appearance
of D-cracking will be at the intersection of transverse and
longitudinal joints in a pavement. At these locations, moisture is often available at the tops and the bottoms of the
vertical joint faces.
In climates cold enough to freeze through the thickness
of the concrete slab, D-cracking usually starts at the bottom
and progresses to the surface (Schwartz 1 987). This is probably due to the greater availability of moisture at the bottoms
of slabs. By the time telltale cracks appear on the surface,
deterioration can extend 1 .5 ft (0.5 m) or more away from
the joint.
The appearance of D-cracking in milder climates is somewhat different because the concrete never freezes all the way
through. D-cracking in these cases often appears as shallow
spalling at joints; closer examination reveals the characteristic cracks parallel to the joint.
4.3.1.3 Vertical construction —Though much less
common, D-cracking can also appear on vertical construction. Construction details, maintenance practices, or both,
that allow the accumulation of moisture against the corners
of walls or columns can contribute to D-cracking if susceptible aggregates were used in the construction. An example
of such maintenance practices at a building would be the
shoveling of snow off a sidewalk and depositing it against a
concrete foundation wall.
4.3.2 Prevention o f D-cracking
4.3.2.1 Role o fmixture proportioning —The primary factor
in a concrete mixture that contributes to the development
of D-cracking is the susceptibility of the coarse aggregate,
whereas the air-void system and the w/cm have little or no
effect (Schwartz 1 987; Missouri Highway and Transportation Department 1 990 ) . Mos t coars e aggregates identifed as
susceptible to D-cracking are sedimentary rocks, although
many sedimentary rocks have not been found to be susceptible to D-cracking. Igneous rocks are generally not considered to be susceptible to D-cracking unless the rocks are
weathered. Weathered rocks would probably be undesirable
for concrete production anyway due to their low strength and
likelihood to break down from handling. Most metamorphic
rocks have not shown D-cracking susceptibility; however,
some partially metamorphosed sedimentary rocks have been
identifed as s usceptible ( Stark 1 976).
The maximum aggregate size is also important in the
development of D-cracking. Numerous studies (Stark and
Klieger 1 973 ; Klieger et al. 1 974; Stark 1 976; Missouri
Highway and Transportation Department 1 990; Almond
and Janssen 1 991 ) have shown that reducing the nominal
18
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
maximum size of the aggregate reduces its susceptibility.
Unfortunately, reducing the nominal maximum aggregate
size can have less-desirable side effects, including increased
paste demand to maintain workability at a given strength
level, increased drying shrinkage potential, and reduced
joint load transfer in pavements.
4.3.2.2 Importance o f aggregate identifcation —Dcracking can require a number of years to fully develop,
during which time much susceptible concrete could be
placed be fore a problem is identifed. This, combined with
natural variability of aggregate sources, leads to the need for
identifcation o f D-cracking-susceptible aggregates be fore
they are used in concrete exposed to moisture and cycles of
freezing and thawing.
D-cracking has been known since the 1 930s (Stark and
Klieger 1 973 ); wide ranges of tests have been developed to try
to identify susceptible aggregates. The most common procedure for identifying susceptible aggregates is ASTM C666/
C666M. Concrete specimens made with the aggregate in question are subjected to repeated cycles of freezing and thawing
in the laboratory and are evaluated in terms of either increase
in length or decrease in dynamic modulus of elasticity.
4.3.2.3 Aggregate benefciation —A variety of techniques
have been proposed to improve the performance of susceptible aggregates. Limiting the nominal maximum size was
discussed in 4.3.2.1 . Schwartz (1 987) summarized other
methods, including coating the aggregates to prevent their
absorption of water, heavy media separation, and blending
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from
durable aggregate with
a nondurable
aggregate
to reduce
its D-cracking susceptibility. He reported that aggregate
size reduction was the most effective method of reducing
D-cracking susceptibility.
4.3.3 Mitigation o f existing D-cracking —A considerable
amount of concrete containing D-cracking-susceptible aggregates has been placed where it is exposed to both moisture
and cycles of freezing and thawing. This is especially true
of concrete pavements. Joint deterioration associated with
D-cracking can signifcantly reduce the service life o f such
pavements. The concrete as little as 1 .5 ft (0.5 m) away from
the joints often shows no deterioration or loss of strength.
A typical concrete pavement can have transverse joints
1 2 ft (4 m) or more apart. With less than 1 .5 ft (0.5 m)
of D-cracked concrete at each end of a slab, most of the
concrete is in good condition. While replacement of the
deteriorated concrete near the joints with a full-depth patch
would seem to be a cost-effective method of extending the
life, D-cracking often appears at the newly created joints
adjacent to the patches in as little as 5 years. Thus, the
D-cracking continues as before, but at twice as many joints
(Janssen and Snyder 1 994).
4.3.3.1 General —Three conditions are necessary for the
development of D-cracking: concrete made with susceptible
aggregate, moisture, and cycles of freezing and thawing. As
the aggregates already in the concrete would be di ffcult or
impossible to render nonsusceptible, D-cracking mitigation
should attempt to either prevent freezing and thawing or
remove the source of moisture.
4.3.3.2 Preventing freezing—Portland-cement concrete
pavements often receive asphalt concrete overlays to improve
the condition of the pavement and extend its life. In climates
that do not get too cold in winter, freezing in a concrete pavement that contains D-cracking-susceptible aggregates could
be prevented by covering it with a suffcient thickness o f
asphaltic concrete. Janssen et al. (1 986) found that the freezing
should be almost completely prevented in the concrete to stop
the progression of D-cracking; merely decreasing the number
of cycles of freezing and thawing with an overlay could actually accelerate the rate of D-cracking due to the increased
potential for moisture migration during slower freezing rates.
More than 6 in. (1 50 mm) of asphalt concrete overlay would
be required to prevent freezing at the surface of the concrete
pavement for central latitudes of the United States. Using
asphalt concrete overlay to prevent freezing in concrete made
with D-cracking-susceptible aggregates is probably not an
effective D-cracking mitigation method for these conditions
(Janssen et al. 1 986; Janssen and Snyder 1 994).
4.3.3.3 Reducing moisture —The use of sealers on the
cut ends of the existing concrete pavement sections (4.3.3)
before placing the patches, could reduce the lateral movement of moisture into the concrete. This could increase the
time before D-cracking appeared in the patched concrete.
This method was attempted in a D-cracked concrete pavement section in Ohio in 1 992 (Janssen and Snyder 1 994).
Though initial laboratory testing indicated that sealer treatment delayed the resumption o f D-cracking, feld moni Standard
Sharing
and ourreappeared
chats after six years
toring showed
thatGroup
the D-cracking
(Janssen 2001 ).
4.3.4 Theories and mechanisms o f D-cracking —Theories of damage to concrete from freezing and thawing have
already been discussed in 4.2.5. With the exception of the
role of air voids in protecting concrete from damage, these
same theories generally apply to D-cracking. This section,
there fore, will concentrate on the characteristics o f specifc
aggregates that make them susceptible, while other aggregates of the same type are not susceptible.
4.3.4.1 Pore size and size distribution —Kaneuji et al.
(1 980) observed qualitative correlations between concrete
durability and pore size distributions of aggregates. At a
constant total pore volume, aggregates with smaller pores
result in a lower resistance to freezing and thawing. For
aggregates with similar predominating pore sizes, a greater
pore volume results in less resistance to freezing and thawing
aggregate. By correlating aggregate service records with
mercury porosimetry studies, Marks and Dubberke (1 982)
found that with one exception, the D-cracking-susceptible
aggregates analyzed exhibited a predominance of pore sizes
of 1 .5 × 1 0 –6 to 8 × 1 0 –6 in. (0.04 to 0.2 μm) whereas aggre gates with good to excellent service records had a majority
of pores that were larger than this.
4.3.4.2 Deicing salt effect—Dubberke and Marks (1 985)
noted a reduced resistance to D-cracking for some aggregates
when pavements containing susceptible aggregates were
exposed to deicing salt. Other aggregates that they examined showed no effect. A possible explanation is a change
in pore structure due to etching of the pore walls by the salt.
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
action has been identifed for aggregates
containing
calcite (Gillott 1 978). This possible deicing salt effect should
be noted when using feld performance records to evaluate
the D-cracking potential of an aggregate. Satisfactory performance in a pavement that never received deicing salt may not
ensure the same performances where deicing salts are used.
S uch
CHAPTER 5—ALKALI-AGGREGATE REACTION
5.1 —Introduction
This chapter presents guidelines for minimizing the risk
of damaging expansion caused by alkali-aggregate reaction (AAR) in concrete construction. This risk may include
potential damage from alkali-carbonate reactive or alkalisilica reactive aggregates. Procedures are discussed for evaluating aggregates and selecting appropriate measures for
controlling expansion are discussed. ACI 221 .1 R provides
more detailed information on types of reaction, reaction
mechanisms, reactive rock types, methods of testing aggregates, and preventive measures.
Alkali- s ilica
reaction
(AS R)
was
frs t detected
in the
late
Stanton 1 940a,b). Since its
f
premature concrete deterioration throughout the world. Since
S tanton’s frs t inves tigations , a wide range o f tes ting proce dures for assessing aggregate reactivity has been developed.
Most recently, effort has focused on developing accurate
testing methods to determine the effectiveness of mitigation
measures. Much focus has been placed on the incorporation
of supplementary cementitious materials (SCMs) as well as
chemical admixtures, namely, lithium salts such as LiNO 3 .
The reliability of these techniques varies and depends on, to
some extent, the nature of the aggregate being tested and the
testing environment.
The
frs t cas e
o f alkali- carbonate
reaction
(AC R)
was
observed in Ontario, Canada, in the late 1 950s (Swenson
1 957). ACR occurs between alkali hydroxides and certain
argillaceous dolomitic limestones. This reaction is characterized by rapid expansion and extensive cracking of the
affected concrete. Structures suffering from ACR generally exhibit deleterious effects within 5 years or less from
initial construction. The only way to avoid ACR is through
selective quarrying to avoid construction with potential
alkali-carbonate reactive aggregates. This type of reaction is
limited to select geographical regions.
Criteria for interpreting test results vary among different
national standards. They also differ among and within states
or provinces, with different limits being adopted by various
local agencies and state or provincial authorities. Methods
for controlling expansion due to AAR also vary regionally. Many s pecifcations do not permit the us e o f reactive
aggregates. When reactive aggregates are used in concrete,
recommended preventive measures include limiting the
alkali content o f the concrete; us ing S CMs s uch as f y as h,
slag cement, silica fume, or natural pozzolans; using lithium
nitrate; or combining these methods.
A his tory o f s atis factory feld per formance may be the
most effective method for evaluating the potential for an
1 9 3 0 s and frs t reported in 1 9 40 (
initial detection, AS R has been identifed as a maj or caus e o
19
aggregate to cause AAR (5.3.1 ). Where such satisfactory
feld per formance can be demonstrated, aggregates may be
accepted for use in concrete without AAR testing, provided
that similar materials are incorporated in the batching
process. An example is concrete with an alkali content less
than or equal to that of the satisfactory concrete in service.
In the abs ence
o f such feld per formance
data,
however,
aggregates should be subjected to suitable laboratory testing
procedures to establish their degree of reactivity. If the
results of such laboratory testing do not indicate a potential
for AAR, aggregates may be used without any precautionary
measures. Even aggregates that demonstrate the potential
for ASR may be used in concrete, provided that suitable
measures are implemented to control the risk of expansion.
Alkali-carbonate reactive aggregates are normally avoided,
however, as it has been proven di ffcult, or economically
unfeasible, to control expansion with such materials.
5.2—Types o f reactions
Two types of AAR have been recognized: ACR and ASR.
Alkali-carbonate reaction is associated with the use of certain
argillaceous dolomitic limes tones . Confrmed cas es o f ACR
have been restricted to a few locations in North America:
mainly in Virginia, Kentucky, Indiana, Iowa, Illinois, and in
Ontario, Canada. Alkali-carbonate reaction involves a reaction between an alkali source and certain calcium-magnesium carbonate rocks (dolomites). Alkali-silica reaction
is distinctly different, and results from a reaction between
alkali hydroxides in the pore solution and certain forms of
reactive silica present in some types of siliceous or carbonate
aggregates. Alkali-silica reaction can occur in limestone
aggregates that contain siliceous components as well. Table
5.2 presents a list of several common reactive rock types and
mineral forms that are susceptible to ASR.
Alkali-silica reaction is far more widespread than ACR,
and is further subdivided into two categories: 1 ) reactions
involving poorly crystalline or metastable silica materials;
and 2) reactions involving certain varieties of quartz. From
an engineering perspective, the main distinctions between
these two categories of ASR involve the time to the onset of
expansion and cracking, and the perceived duration of the
reaction in the feld. Reactions involving s uch s ilica mate rials, which are sometimes referred to as classical ASR,
are characterized by a relatively short time to the onset of
cracking where cracking usually occurs within 5 to 1 0 years,
whereas the manifestation of reactions involving quartz
minerals usually takes much longer, although the reaction
may continue for many decades.
5. 2 . 1 ACR background —The ACR occurs between alkali
hydroxides and certain argillaceous dolomitic limestones;
thes e dolomites are characterized by a matrix o f fne calcite
and clay minerals with scattered dolomite rhombohedra. The
reaction is manifested in the rapid expansion and extensive
cracking of concrete; structures affected by ACR usually
show cracking within 5 years. Although there is a lack of
consensus regarding the precise mechanisms involved, it
is generally agreed that the reaction is accompanied by the
dedolomitization process, as follows
20
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
Table 5.2—Some examples o f rock types and
minerals susceptible to ASR
Reactive rocks
Reactive minerals
Shale
Opal
Sandstone
Tridymite
Silicifed carbonate rock
Cristobalite
Chert
Volcanic glass
Flint
Cryptocrystalline (or microcystalline)
quartz
Quartzite
Strained quartz
Quartz-arenite
Gneiss
Argillite
Granite
Greywacke
Siltstone
Arenite
Arkose
Hornfels
Ca· Mg(CO 3 ) 2 + 2ROH → CaCO 3 + Mg(OH) 2 + R2 CO 3
(5.2.1 a)
(dolomite + alkali hydroxide → calcite + brucite + alkali carbonate)
b), ASR has been observed as a cause of premature concrete
deterioration throughout the world. Although the factors that
lead to deleterious ASR are commonly agreed on, the mechanism by which the alkali-silica gel causes expansion and
subsequent cracking in concrete is not yet entirely understood by researchers in the feld.
Alkali-silica reaction is a chemical reaction that is the
result of hydroxyl ions attacking siliceous species in certain
aggregates. The attack liberates silica, which then combines
with alkalis (Na + and K+) and with lesser amounts of calcium
(Ca++) that are present in the concrete pore solution to maintain charge balance. The resulting alkali-silica gel then
absorbs water and expands, which may result in cracking
of the aggregates, the cement paste, and ultimately the
concrete matrix. For the ASR to cause damage in concrete,
it is widely accepted that three components are necessary:
suffcient alkali, reactive silica, and adequate moisture.
5.2.2.1 A lkalis —The alkalis (Na + and K+) are typically
supplied by portland cement. However, SCMs; chemical
admixtures; and external sources such as seawater, deicing
salts, and anti-icing chemicals can also contribute to the
alkalinity of the pore solution. Certain aggregate species,
particularly those containing feldspars, may also release
alkalis to the pore solution (Bérubé et al. 2002). The amount
of alkali in cement is usually expressed as the sodium oxide
equivalency, written Na2 O eq.
Equation (5.2.2.1 ) is used to determine the sodium oxide
equivalency in the portland cement
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where R represents K or Na. Because this reaction results in
a reduction in solid volume, however, the expansion must
be attributed to an alternative mechanism. Several theories
have been proposed to explain the expansion mechanism
(Swenson and Gillott 1 964; Tang et al. 1 987; Fournier and
Bérubé 2000), which include:
a) Hydraulic pressures caused by the migration of water
molecules and alkali ions into the restricted spaces of the
calcite/clay matrix around the dolomite rhombs
b) Adsorption of alkali ions and water molecules on the
surfaces of the active clay minerals scattered around the
dolomite grains
c) Growth and rearrangement of the products of dedolomitization (brucite and calcite)
The alkali carbonate produced in the dedolomitization
reaction may react with calcium hydroxide (CH) in the
cement paste as follows
R2 CO 3 + Ca(OH) 2 → CaCO 3 + 2ROH
(5.2.1 b)
thereby regenerating alkalis for further reaction. Thus,
provided there is su ffcient alkali available to initiate the reac tion, the process may continue independently of the amount
of alkalis available in the concrete. This could explain why
low-alkali cements are not effective in controlling damaging
reaction in some instances (Thomas and Folliard 2007).
5.2.2 A SR background —Alkali-silica reaction has been
a known cause of concrete deterioration for more than 70
years. Since its discovery in the late 1 930s (Stanton 1 940a,
Na2 O eq = Na2 O + 0.658K 2 O
(5.2.2.1 )
where Na2 O eq is the total sodium oxide equivalent, in percent
by mass; Na2 O is sodium oxide content, in percent by mass;
and K2 O is potassium oxide content, in percent by mass.
The concentration of alkalis in portland cement generally
ranges from 0.2 to 1 .3 percent Na2 O eq, which is relatively
low in comparison to other compounds and oxides. Initial
research on ASR proposed that expansion due to ASR was
unlikely to occur when the percentages of alkalis in the
cement fell below 0.6 percent Na2 O eq (Stanton 1 940b). This
approach has since been used as a mitigation option to limit
ASR in new concrete. Reducing the percent contribution
of alkalis from portland cement, however, does not effectively mitigate ASR for all reactive aggregate types because
it does not limit the total alkali content of the concrete or
alkali loading; within the pore solution of concrete, the
alkalis dissociate in solution, leaving K+ and Na+, which
must then be balanced by an equivalent concentration of
hydroxyl ions (OH –) to maintain charge equilibrium. The
increased concentration of dissociated alkalis in the concrete
pore solution effectively increases the concentration of
hydroxyl ions, which in turn increases the pH in the pore
solution. As referenced in 5.2.2.4, this OH –-induced increase
in pH, if high enough, leads to the initial breakdown of the
reactive silica in the aggregate, resulting in the formation
of alkali-silica gel. Diamond (1 983a) reports that the OH –
ion threshold concentration for ASR is unlikely to be less
than 0.25M and Kollek et al. (1 986) suggest a threshold of
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
0.20M; this equates to a pH threshold of 1 3.2 to 1 3.3 for
ASR to occur.
5.2.2.2 Reactive silica —The degree of reactivity of aggregates depends on a number of factors, including the mineralogy of the aggregate, and the crystallinity and solubility
of the silica. Reactive silica is amorphous or disordered
silica found in certain aggregates. This poorly crystalline
silica dissolves more readily in the alkaline pore solution of
concrete than do well-crystallized or dense forms of silica.
Because of its increased solubility, amorphous silica is more
susceptible to ASR. Common reactive minerals susceptible to ASR include microcrystalline, cryptocrystalline,
and strained quartz; cristobalite; tridymite; opal; chert; and
volcanic glass such as obsidian.
5.2.2.3 Adequate moisture —The third and fnal compo nent necessary for ASR to occur is adequate moisture, which
is one of the key components in the expansion of the gel.
Water is found within the pore solution of concrete and is
also introduced from external sources. A minimum relative
humidity of 80 percent is required to provide enough moisture to drive the expansion of the alkali-silica gel and sustain
the reaction (Pedneault 1 996).
5.2.2.4 Mechanism o f gel formation —The term “ASR” is
somewhat misleading, as the initial reaction occurs between
the reactive siliceous aggregate and the hydroxyl (OH –) ions
and not the alkalis of the pore solution. The high concentration of the hydroxyl ions in the pore solution is equal to that
of the alkali cations to maintain charge equilibrium. This
high OH – concentration is what results in a high pH, which,
in turn, leads to the initial breakdown of the reactive silica
in the aggregate. When poorly-crystalline hydrous silica is
exposed to a strong alkaline solution, there is an acid-base
reaction between the hydroxyl ions in solution and the acidic
silanol (Si-OH) groups (Dent Glasser and Kataoka 1 981 ) as
follows
≡Si-OH + OH – → ≡Si-O- + H 2 O
(5.2.2.4a)
As further hydroxyl ions penetrate the structure, some of
the siloxane linkages (Si-O-Si) are also attacked as follows
(Dent Glasser and Kataoka 1 981 )
≡Si-O-Si≡ + OH – → 2 ≡Si-O- + H 2 O (5.2.2.4b)
The negative charges on the terminal oxygen atoms are
balanced by alkali cations (Na+ and K+) that simultaneously
diffuse into the structure. The disruption of siloxane bridges
weakens the structure and, provided suffcient reserves o f
alkali hydroxide are available, the process continues to
produce an alkaline silicate solution. The extent or rate of
dissolution is controlled by the alkalinity of the solution and
the structure of the silica.
5.2.2.5 Role o f calcium —Bleszynski and Thomas (1 998)
and Thomas (2006a) concluded that signifcant expansion
only occurs when an adequate supply of calcium is available as CH. In systems with abundant alkali hydroxides
and reactive silica but no CH, silica dissolved and remained
in solution. Although the precise role calcium plays in gel
21
expansion remains unclear, a series of mechanisms has been
proposed:
a) Calcium can replace alkalis in the reaction product,
regenerating alkalis for further reaction (alkali recycling)
(Thomas 2006b; Hansen 1 944)
b) Calcium hydroxide may act as a buffer maintaining a
high level of OH – in solution (Wang and Gillott 1 991 )
c) High calcium concentrations in the pore solution prevent
the diffusion of silica away from reacting aggregate particles
(Bleszynski and Thomas 1 998 ; Chatterji 1 979; Chatterji and
Clausson-Kaas 1 984)
d) If calcium is not available, reactive silica may merely
dissolve in alkali hydroxide solution without causing damage
(Thomas 2006a; Diamond 1 989)
e) The formation of calcium-rich gels is necessary to
cause expansion either directly or through the formation
of a semi-permeable membrane around reactive aggregate
particles (Thomas 2006a; Thomas et al. 1 991 ; Bleszynski
and Thomas 1 998)
Promoting the formation of C-S-H at the expense of CH
(for example, through the use of pozzolans) may result in
successful mitigation of expansion due to ASR.
5.2.3 Mechanism o f gel expansion —Although the mechanisms behind the formation of the gel are well understood,
the actual mechanism for expansion of gel remains uncertain. Four main theories have emerged over the past 70 years
to explain the mechanism of gel expansion, all maintaining
that water is the main component driving the process. The
four theories of expansion include the double-layer, osmotic
pressure, CSH-shell, and the calcium/alkali exchange theory.
The osmotic pressure theory speculates that the cement
paste surrounding the reactive aggregates acts as a semipermeable membrane, preventing the presence of large silicate ions while allowing the water and alkali hydroxides to
diffuse through. Under these conditions, the alkali silicate
formed on the surface of the aggregate particle draws solution through the cement paste, resulting in continued formation of the alkali silica gel. As the gel continues to swell, an
osmotic pressure cell is formed and increasing hydrostatic
pressure is applied to the cement paste, eventually resulting
in cracking (Hansen 1 944).
The C-S-H shell theory considers the effect of calcium
on the durability of concrete. This theory hypothesizes
that in the presence of CH, alkali ions from alkali salts and
hydroxyl ions from Ca(OH) 2 enter the reactive silica aggregate grains, leaving calcium and anions in the pore solution.
The penetration of the solvated hydroxyl and alkali ions
causes the Si-O-Si bonds of the reactive aggregate to break
apart, opening the grains for further penetration of ions,
and permitting the release of some SiO 2 into the pore solution. As the solvated hydroxyl and alkali ions infltrate the
aggregate grains, calcium, hydroxyl, and water molecules
also migrate into the reactive siliceous material. When high
concentrations of Ca(OH) 2 and alkali salts are present in the
pore solution, only a limited amount of SiO 2 can diffuse out,
while additional materials penetrate into the aggregate structure. This imbalance results in an expansive force within
the aggregate grain. If, however, the Ca(OH) 2 and alkali
22
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
salt concentrations are low, the rate of penetration of the
hydroxyl and alkali ions is also low, while the migration of
SiO 2 is increased, thus resulting in a lower expansive force
(Chatterji et al. 1 987).
The calcium/alkali exchange theory also considers calcium
to be an essential component in the expansion of alkali-silica
gel. This theory hypothesizes that the gel absorbs Ca2+ ions
and, in turn, exchanges them for alkali ions, which then react
with the pore solution to create additional alkali-silica gel.
As additional gel forms , it flls more s pace and eventually
causes cracking. The ion and water uptake is governed by
the temperature and moisture conditions of the material and
thus is coupled to physical transport (Rotter 1 995 ).
The double-layer theory suggests that the expansion of
the gel is caused by swelling due to electrical double-layer
repulsive forces. When a liquid and a solid come into contact,
the surface of the solid carries excess charge, which electrifes the inter face.
This excess
charge alters
the properties
of both the solid and liquid materials. Alkali-silica reaction
involves the interaction of the highly-charged silica aggregate surface with the alkaline pore solution of the concrete.
This is the reaction that leads to the breakdown of the silica
and the formation of alkali-silica gel. Within the gel, negatively charged solid silica particles attract positively charged
cations that thus bind to form a rigid layer around the solid
particle. Surrounding this rigid layer, a diffuse layer is formed
that comprises more cations and anions found within the gel.
The electrical double layer, therefore, is composed of sodium,
standards
from
potassium, and calciumGet
ions,more
whichFREE
surround
the negatively
charged silica surface. Once this double layer has been established, it imbibes water and swells. As water is introduced into
the layer, electrostatic forces predominate and particles are
pushed apart as the gel expands (Prezzi et al. 1 997).
While it is commonly understood that water is the primary
driving force for expansion of alkali-silica gel, no single
theory about the mechanism of gel expansion is widely
agreed upon, and none appears to have completely and accurately explained the mechanism. For example, none of the
four main theories discussed previously considers the potentially crucial role that short-range forces might play in expansion. Neither the osmotic pressure theory nor double-layer
theory considers the potential effect of calcium on expansion. Although at this time it remains unknown exactly how
alkali-silica gel expands in concrete, portions of the theories
described previously may together explain the mechanism
or lead to a more complete explanation.
5.3—Evaluating aggregates for potential alkaliaggregate reactivity
5.3.1 Field performance —A his tory o f satis factory feld
performance is possibly the best method for evaluating
the potential for an aggregate to cause AAR. A number of
factors have to be cons idered when analyzing feld performance data. These include:
a) The cement content of the concrete and the alkali
content of the cement should be the same or higher in the
feld concrete as propos ed in the new s tructure.
b) The concrete examined should be at least 1 0 years old
and preferably more than 1 5 to 20 years old.
c) The expos ure conditions o f the feld concrete s hould be
at least as severe as those in the proposed structure.
d) In the absence of documentation conclusively demonstrating that the aggregate to be used in the proposed structure is s u ffciently s imilar to the feld structure under inves tigation, a petrographic examination should be conducted to
make that determination.
e) The possibility that SCMs, lithium-based admixtures, or
both, were us ed in the feld s tructure s hould be cons idered.
f) P rovided that satis factory feld performance can be
demonstrated, the aggregate can be used in concrete,
following the prior-listed guidance, with no further testing
for AAR.
5.3.2 Petrographic examination (ASTM C295/C295M) —
A
petrographic
examination
s hould
be
the
frst
s tep
in
assessing the suitability of a particular aggregate source
for use in concrete construction. Petrography is a powerful
tool that yields a wide range of information regarding the
physical, chemical, and mineralogical characteristics of an
aggregate, including the presence of rocks or mineral phases
that are known to cause deleterious reaction in concrete.
In some cases, a petrographic examination may produce
s u ffcient evidence to rej ect an aggregate on the basis o f
potential alkali reactivity or require that suitable preventive measures be explored and implemented. Generally, the
examination cannot predict whether the type and distribuStandard
Sharing
Grouppresent
and our
tion of reactive
minerals
will chats
cause damaging expansion in concrete, and further laboratory testing is usually
required. Results of petrographic examination could form
the basis for directing the laboratory test program in terms
of selecting the type and sequence of tests and any relevant
evaluation criteria.
The reliability of petrographic examination for screening
aggregates for potential reactivity is strongly dependent
on the skill and experience of the individual petrographer.
There have been cases where aggregates that were accepted
for use on the basis of results of petrographic examination
have been later implicated in AAR. This is not necessarily
the res ult o f incorrect material clas sifcation, but more likely
is due to a failure to recognize certain minerals as potentially
reactive (Rogers and Hooton 1 991 ). Furthermore, the reactive cons tituents o f s ome rocks may not be readily identifed
by optical microscopy.
5.3.3 Laboratory tests to identi fy alkali-silica reactive
aggregates —Many test methods have been developed for
identifying alkali-silica reactive aggregates. These methods
vary in terms of testing environment, duration of test, and
reliability of results. A comprehensive review is provided by
Thomas et al. (2006). Generally, the tests of longer duration such as the concrete prism test (5.3.3.4) produce more
reliable results than shorter-duration, highly aggressive tests
such as the accelerated mortar bar test (Thomas et al. 2006).
In this section, a brief description is provided of test methods
to detect ASR in aggregates.
5.3.3.1 Mortar bar test (ASTM C227) —This test method
was originally developed to assess the potential alkali-silica
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
reactive of cement-aggregate combinations and is not recommended for evaluating ACR aggregates. Mortar bars (aggregate/cement = 2.25) are monitored for expansion under
storage conditions at 1 00°F (38°C) in storage containers
designed to maintain high humidity. For testing, in ASTM
C227 coarse aggregates are crushed to a fne aggregate size
fraction and both coarse and fne aggregates follow specifc
gradation requirements set forth in ASTM C227. The alkali
content o f the cement is not specifed, but the cement should
be selected to have the highest alkali content representative
of the cement generally intended for use with the aggregate.
The standard specifcation for concrete aggregates
(ASTM C33/C33M ) regards as reactive those cementaggregate combinations that have an expansion greater than
0.05 percent at 3 months or 0.1 0 percent at 6 months. More
stringent limits have been applied by certain agencies. For
example, the U.S. Bureau o f Reclamation requires specifc
combinations to expand by less than 0.05 percent at 6
months, or 0.1 0 percent at 1 2 months.
There are a number of limitations associated with this
test. For example, it is di ffcult to attain high humidity in the
containers without leaching the alkalis from the mortar bars.
Given the relatively small size of these mortar bars (1 x 1 in.
[25 x 25 mm] cross section), leaching may be severe, especially during early testing period. Alkali leaching can lead to
an underestimation of the expansion of certain combinations
of cement and aggregate, especially if the reactive component of the aggregate reacts relatively slowly. This has been
observed in argillites and greywackes in which the reactive component is microcrystalline quartz (Grattan-Bellew
1 989; Rogers and Hooton 1 991 ). The consequences of alkali
leaching may be reduced by storing bars over water without
wicks and by raising the equivalent alkali content or Na2 O eq
(note: Na2 O eq = Na2 O + 0.658 K2 O) of the cement to 1 .0
percent, but expansions will still be less than 0.1 0 percent at
6 months with some reactive aggregates. Due to these limitations, ASTM C227 is not recommended for identifying
alkali-silica reactivity of aggregates or for cement-aggregate
combinations.
5.3.3.2 Quick chemical method (ASTM C289) —In this
test, a sample of the aggregate, crushed to pass a No. 50
(300 μm) sieve and retained on a No. 1 00 (1 50 μm) sieve,
is immersed in 1 molar NaOH solution for 24 hours, and
the resulting solution is analyzed. The amount of silica
dissolved and the reduction in alkalinity of the host solution
are plotted on a graph with zones that classify the aggregate
as innocuous, potentially reactive, or deleterious.
The reliability of this test in detecting aggregate reactivity is poor for a number of reasons. Other mineral phases
present in the aggregate may reduce the dissolved silica by
precipitation (Bérubé and Fournier 1 992a). Furthermore,
reactive phases may be lost during crushing and sieving.
Thus, reactive aggregates may appear innocuous based on
the test results. In contrast, the high surface area and temperature used in this test dissolve some siliceous mineral phases
that are stable under the conditions that prevail in concrete.
This results in aggregates with good feld performance being
23
classifed as deleterious (Bérubé and Fournier 1 993 ). Note
that ASTM C289 was withdrawn by ASTM in 201 6.
5.3.3.3 Accelerated mortar bar test (ASTM C1260) —This
test is essentially the same as that developed by Oberholster and Davies (1 986). In this method, the length change of
mortar bars (measuring nominally 1 x 1 x 1 1 .25 in. [25 x 25 x
285 mm]) stored in 1 molar NaOH solution at 1 76°F (80°C)
is monitored for 1 4 days. For testing, in ASTM C1 260, coarse
aggregates are crushed to a fne aggregate size fraction and
both coarse and fne aggregates follow specifc gradation
requirements set forth in ASTM C1 260. Some agencies
specify that length change be measured over 28 days in 1
molar NaOH. The expansions obtained in this rapid test are
generally comparable to or higher than those obtained by
ASTM C227 (1 00°F [38°C] at 1 00 percent humidity) after
1 year (Oberholster and Davies 1 986; Hooton and Rogers
1 989). The test has been successfully used to identify alkalisilica reactive aggregates from across Canada (GrattanBellew 1 989; Hooton and Rogers 1 989, 1 992; Bérubé and
Fournier 1 992b; Durand et al. 1 990; Hooton 1 991 ) and the
United States (Stark et al. 1 993 ), but is not considered suitable for evaluating reactive alkali-carbonate rocks.
Interpretation of results is not simple, and various expansion criteria have been suggested (Thomas et al. 2007).
Bérubé and Fournier (1 992b) proposed a limit of 0.1 0
percent expansion after 1 4 days in 1 M NaOH for quarried
silicate and siliceous carbonate aggregates, and a limit of
0.20 percent for natural sands and gravels. Many aggregates with satis factory feld performance, however, produce
expansions in excess of 0.25 percent in this test (Bérubé and
Fournier 1 992b). Consequently, aggregates should not be
rejected on the basis of this test unless petrographic examination confrms that the material is similar to known delete riously reactive aggregates.
Recent research has shown that certain aggregate types
may pass this test, having expansions less than 0.1 0 percent
at 1 4 or 28 days, yet cause deleterious expansion in ASTM
C1 293 and in the feld ( Ideker et al. 201 2). In summary, this
test should be used with caution, owing to the many potential discrepancies between the performance of aggregates in
this test and in the feld (Thomas et al. 2007).
5.3.3.4 Concrete prism test (ASTM C1293) —This test is
considered the most reliable for correctly identifying alkalisilica reactive aggregates (Thomas et al. 2006). A concrete
mixture is proportioned with a cement content of 708 ± 1 7
lb/yd 3 (420 ± 1 0 kg/m 3 ) using a portland cement with an
equivalent alkali content (Na2 O e) of 0.90 ± 0.1 0 percent.
Sodium hydroxide is then added to the mixing water to
provide a total alkali loading of the concrete of 8.85 lb/yd 3
(5.25 kg/m 3 ). This high alkali loading is necessary to induce
expansion of slowly reactive rocks such as greywacke and
argillites (Magni et al. 1 987). Prisms are stored at 1 00°F
(38°C) and the expansion is monitored for at least 1 year.
An expansion limit of 0.04 percent at 1 year is currently
specifed in both Canada and France (Thomas et al. 1 997)
and is recommended in ASTM C1 293. This expansion correlates approximately to the point where cracking and signs of
distress are frst observed on the prisms. It also relates well
24
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
f
CSA A23.2-27A and CSA A23.2-28A
carry a caveat stating that lower expansion limits may be
appropriate for assessing aggregates for use in critical structures, such as nuclear containment, or dimensionally sensitive structures, such as hydraulic dams, where small expansions may result in relatively large movements.
This test method has distinct advantages over the mortar
bar tests (ASTM C227; ASTM C1 260), as coarse aggregates
can be tested without crushing to sand sizes. Furthermore,
the larger test specimen reduces the effect of alkali leaching.
The 1 2-month duration is necessary unless the temperature
is raised or other changes are made to accelerate expansion;
however, further acceleration of the test may have undesirable side effects (Ideker et al. 201 2).
5.3.3.5 Accelerated concrete prism test (RILEM AAR 4 –
proposed) —Due to the duration of the concrete prism test,
there has been a push since the early 1 990s to accelerate the test
by increasing the temperature from the standard 1 00 to 1 40°F
(38 to 60°C) (Ranc and Debray 1 992; Murdock and Blanchette
1 994; DeGrosbois and Fontaine 2000; Touma et al. 2001 ).
Unfortunately, simply raising the temperature and identifying expansion limits was not enough to produce a consistent and reliable test method. Fournier et al. (2004) identifed factors
leading to the variability
in the accelerated
test, including increased leaching of alkalis, reduced pore
solution pH, increased mass loss, and a potential concern
related to s election o f the nonreactive fne aggregate; they
also proposed expansion limits. Further work by Ideker et
Getprofound
more FREE
standards
from
al. (201 0) made clear the
effect of
the selection
of
the nonreactive fne aggregate on expans ions in the acceler ated test: expansions were reduced by as much as 50 percent
compared to the 1 -year expansions obtained with the Spratt
coarse aggregate.
The accelerated version of the concrete prism test results
in a reduction in expansion at 3 months as compared to
the 1 -year expansion obtained in ASTM C1 293 (Ideker et
al. 201 0). Thus, the accelerated concrete prism test is not
recommended for assessing aggregate reactivity or for determining the e ffcacy o f mitigation meas ures .
5.3.3.6 Chinese accelerated concrete microbar test (for
alkali-silica reactive aggregates) —A new test method,
now commonly referred to as the concrete microbar test
(formerly the Chinese accelerated mortar bar method), was
introduced by Xu et al. (1 998 , 2000) to capture reactivity of
alkali-carbonate reactive rocks. In this test method, mortar
bars measuring 1 .58 x 1 .58 x 6.30 in. (40 x 40 x 1 60 mm) are
cast, cured for 24 hours, and then soaked in water at 1 76°F
(80°C) for an additional 24 hours. The bars are then moved
into a 1 N NaOH at 1 76°F (80°C) and length change is monitored for 28 days. Aggregates are graded (or crushed where
appropriate) to produce to a size fraction in the range of 0.2
to feld per ormance.
to
0. 4
in.
(5
to
10
mm)
and cas t into
mortar bars
at a fxed
cement-aggregate ratio and a w/cm of 0.33. Since that time,
promising results have also been obtained for the detection
of ASR across a wider range of rock types (Lu et al. 2008 ;
East 2007). In particular, the ability of this test to detect
deleterious ASR in coarse aggregates is advantageous where
reactive phases are removed due to crushing and processing
in other accelerated methods (ASTM C227; ASTM C1 260).
This accelerated test method is gaining momentum, and
other res earchers have s hown its e ffcacy in testing poten tially alkali-silica reactive coarse aggregates. This method
may serve as a complementary test method, especially when
ASTM C1 260 produces a false negative. Lu et al. (2008)
have recommended it as a universal accelerated test for
alkali-aggregate reactivity.
5.3.4 Laboratory tests to identi fy reactive alkali-carbonate
rock aggregates
5.3.4.1 Rock cylinder method (ASTM C586) —In
this
method, cylinders (or prisms) cut from the rock are immersed
in a solution of 1 M NaOH at room temperature (after having
attained dimensional stability in distilled water) and the
expansion is monitored for at least 1 month. Expansions in
excess of 0.1 0 percent at 1 month are generally taken to indicate a potentially deleterious chemical reaction between the
alkalis and the rock. This test does not provide an indication
of the potential for expansion in concrete, and further testing
of the aggregate in concrete (ASTM C11 05 ) is recommended
if the rock cylinder expansion exceeds 0.1 0 percent.
5.3.4.2 Chemical composition (CSA A23.2-26A ) —The
determination of potential ACR by chemical composition involves analysis for CaO, MgO, and Al 2 O 3 (CSA
A23.2-26A). Limestones or dolomites with a composition
outside of the range indicated as potentially alkali-carbonate
reactive in Fig. 6 of CSA A23.2-26A require further testing
for ASR. Potentially reactive dolomitic limestones plot in
Standard
Sharing
Group area
andofour
chats
the potentially
expansive
a CaO/MgO-versus-Al
2O 3
plot, and such aggregates should be tested by ASTM C1 1 05.
This tes t has helped to remove some o f the di ffculty in
identifying reactive dolomitic limestones by petrographic
examination.
5.3.4.3 Concrete prism test (ASTM C1105) —This test is
similar to ASTM C1 293 used to assess ASR (5.3.3.4), except
for differences in storage temperature and the alkali content of
the concrete. ASTM C11 05 requires the testing to be carried
out using a specifc
concrete
mixture,
with specimens
stored
at 73°F (23°C). Potentially deleterious reactivity is indicated
if the expansion exceeds 0.01 5 percent at 3 months, 0.025
percent at 6 months, or 0.030 percent at 1 year.
Users of this test should recognize that the test yields information
about
the
s pecifc
cement- aggregate
combination
f s ignifcant expans ion in this
test does not necessarily indicate that the aggregate is nonreactive. For instance, deleterious expansion may occur if the
aggregate is used in concrete with a higher alkali content.
CSA A23.2-1 4A requires potentially alkali-carbonate reactive rocks to be tested in concrete prisms under the same
conditions as those used for ASR (that is, with 8.85 lb/yd 3
[5.25 kg/m 3 ] Na2 O e and stored at 1 00°F [38°C]). Furthermore, the same expansion criteria are applied; aggregates
are deemed to be reactive if the expansion exceeds 0.040
percent at 1 2 months, with no criteria for earlier test results.
CSA A23.2-1 4A is aimed at establishing the reactivity of
the rock and not assessing the performance of a particular
cement-aggregate combination.
tes ted,
and that the
abs ence
o
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
5.3.4.4 Chinese accelerated concrete microbar test (for
alkali-carbonate reactive aggregates) —The Chinese accelerated concrete microbar test was created to capture reactivity of alkali-carbonate rocks (Lu et al. 2008 ; Xu et al.
1 998 , 2000; Sommer et al. 2005 ). Further work by Lu et
al. (2008) has shown the ability of this test to detect alkalicarbonate reactive aggregates across a range of aggregate
gradations (0.2 to 0.4 in. [5.0 to 1 0.0 mm] particles and 0.1
to 0.2 in. [2.5 to 5.0 mm] particles). The advantage to using a
0.1 to 0.2 in. (2.5 to 5.0 mm) particle size is that the test can
be used to test a single aggregate sample for both ASR and
ACR. Further, petrographic examination would be needed to
distinguish the type of reaction present so that appropriate
mitigation measures could be followed. This test shows
promise for detecting alkali-carbonate reactive aggregates in
a relatively short time of 4 weeks, while providing more reliable results than the accelerated concrete microbar test. This
test is an attractive alternative to the longer-duration ASTM
C11 05 test for detecting ACR.
5.4—Preventive measures
There are a number of preventive measures that can be used
to minimize the risk of damage due to ASR, which include:
a) Using nonreactive aggregates
b) Limiting the alkali content of the concrete
c) Incorporating SCMs
d) Using chemical admixtures, namely, lithium compo unds
These approaches are discussed in the following sections.
For alkali-carbonate rock reactive aggregates, avoidance or
reduction in proportion to the reactive phases is the only
recommended practice. The other methods listed, though
proven effective with alkali-silica reactive aggregates, are
typically not a remedy for ACR (Rogers and Hooton 1 992).
5.4.1 Use o f nonreactive aggregate —This approach
is perhaps the most obvious and certain way of avoiding
damaging reaction in concrete structures. Nonreactive aggregates are not available in many locations, and
importing nonreactive material may not be economically
viable. Furthermore, AAR has occurred in a number of cases
where prior testing of the aggregates indicated they were
not deleteriously reactive. Methods of testing aggregates for
reactivity have increased in severity, and acceptance criteria
have become more s tringent to re
f
ect the increasing number
f aggregates implicated in feld cas es o f AAR. Adoption
of existing testing practices, however, does not guarantee
that aggregates will give satisfactory performance in every
situation. Consequently, even if aggregates are found not to
be deleteriously reactive, further precautions are frequently
taken as circumstances demand. Such circumstances may
include prestigious (or critical) structures, aggressive environments such as external source of alkalis like seawater or
deicing salts, high cement contents, or extended service life.
5.4.2 Limiting alkali content o f concrete —Stanton’s
(1 940a,b) work on AAR indicated that expansive reaction
is unlikely to occur when the alkali content of the cement
is below 0.60 percent Na2 O e. This value has become the
accepted maximum limit for cement to be used with reactive aggregates in the United States, and appears in ASTM
o
25
C1 50/C1 50M as an optional limit. This criterion, however,
takes no account of the cement content of the concrete that,
together with the cement alkali content, governs the total
alkali content of concrete and is considered a more accurate
index o f potential reactivity.
S ome national s pecifcations
recognize this fact by specifying a maximum alkali content
in the concrete; this limit was reported (Nixon and Sims
1 992) to range from 4.21 to 7.58 lb/yd 3 (2.5 to 4.5 kg/m 3 )
Na2 O e. In some countries, the limit may vary depending on
the reactivity of the aggregate (Oberholster 1 994). In CSA
A23.2-27A, the limit ranges between 3.03 to 5.05 lb/yd 3 (1 .2
and 3.0 kg/m 3 ) Na2 O e. A similar range has been adopted in
Thomas et al. (2008a) and in AASHTO PP065 .
The use of low-alkali cement and limitation of the alkali
content in concrete is not a s u ffcient s a feguard in all cas es .
Stark (1 980) reported damaging AAR in highway structures
constructed using cements with alkalis in the range 0.45 to
0.57 percent Na2 O e. The reactivity of certain aggregates
with low- alkali cements was confrmed in laboratory mortar
bar expansion tests. Lane (1 987) reported that some aggregates, classed as innocuous after 6 or 1 2 months in ASTM
C227 with low-alkali cement (0.54 percent Na2 O e), showed
delayed expansion and cracking after longer periods.
Thomas (1 996) reported evidence of ASR in a number of
hydraulic structures with alkali contents below 4.0 lb/yd 3
(2.4 kg/m 3 ) Na2 O e.
Aggregates that are not normally reactive when used in
concrete with low-alkali cement may be deleteriously reactive in concrete of higher alkali content. This may occur
through alkali concentration caused by drying gradients,
alkali release from aggregates, or the ingress of alkalis from
external sources such as deicing salts or seawater. Stark
(1 978) reported increases in soluble alkali from 1 .85 to 6.07
lb/yd 3 (1 .1 to 3.6 kg/m 3 ) Na2 O e close to the surface of some
highway structures. Migration of alkalis due to moisture,
temperature, and electrical gradients has also been demonstrated in laboratory studies (Nixon et al. 1 979; Xu and
Hooton 1 993 ).
There are many aggregates containing alkalis that may
be leached out into the concrete pore solution, thereby
increasing the risk of AAR (Stark 1 980; Stark and Bhatty
1 986; Way and Cole 1 982; van Aardt and Visser 1 977;
Thomas et al. 1 991 ; Bérubé et al. 2002). Stark and Bhatty
(1 986) reported that in extreme circumstances, some aggregates release alkalis equivalent to 1 0 percent of the portland
cement content.
Alkalis may penetrate concrete from external sources such
as brackish water, sulfate-bearing groundwater, seawater,
or deicing salts. Nixon et al. (1 987) showed that seawater,
or NaCl solutions, present in the mixing water elevates
the hydroxyl-ion concentration and increases the amount
of expansion of concrete. Oberholster (1 992) showed that
the expansion of large concrete blocks exposed to saltwater
spray may be doubled compared with the same blocks
exposed to tapwater spray. In addition, studies in Denmark
(Chatterji et al. 1 987) have shown that exposure to NaCl
solution and other alkali salts can cause considerable expansion and cracking in concrete.
26
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
5.4.3 Use o f SCMs —In his second published article on
ASR, Stanton (1 940b) reported that expansion due to the
reaction could be reduced by the use of pozzolanic cement
containing fnely ground shale or by replacement o f 25
percent of high-alkali portland cement with pumicite. Subsequent tests by Stanton (1 950) confrmed the benefcial e ffect
of a wide range of natural pozzolans and demonstrated that
partially replacing portland cement with a suffcient quan tity of pozzolan (pumicite or calcined shale) eliminated
deleterious expansion, whereas replacement with similar
quantities of ground quartz (Ottawa) sand did not, indicating
that the benefcial action o f the pozzolan extended beyond
merely diluting the cement alkalis. The frst major use o f
a natural pozzolan to control ASR dates back to the 1 940s
when calcined siliceous shale, also called Puente shale, was
used in the Davis Dam, together with low-alkali cement
(Gilliland and Moran 1 949). In the early 1 950s, various
studies (Cox et al. 1 950; Barona de la O 1 951 ; Buck et al.
1 953 ) showed that other SCMs (namely, f y ash and slag)
were also effective in reducing expansion. Later research
showed that silica fume was highly e ffcacious in control ling ASR with levels o f 1 0 percent or less being suffcient
to suppress damaging expansion in mortar bars with reactive aggregate from Iceland (Asgeirsson and Gudmundsson
1 979) and South Africa (Oberholster and Westra 1 981 ).
All SCMs contain alkalis and some, like f y ash, may
contain substantially more than the portland cement they
replace. This has led to considerable controversy in the
standards
from
past regarding whetherGet
the more
alkalis FREE
in SCMs
are potentially
available for reaction and how they should be treated when
calculating the alkali content of the concrete. Recently,
however, it is generally accepted that the main mechanism
by which SCMs reduce the potential for damaging reaction is by reducing the availability of alkali in the concrete
pore solution. Alkalis released by portland cement and SCM
might be available in one of three forms: dissolved in the
pore solution, bound by the products o f hydration, or fxed
by the products of alkali-silica gel. In systems free of reactive aggregate, the partition of the alkalis between the pore
solution and the hydrates is largely a function of the binder
composition (Thomas 201 1 ).
There is a large body of data in the literature that shows
how various SCMs affect the composition of the pore solution extracted from hydrated cement pastes (Longuet 1 976;
Diamond and López-Flores 1 981 ; Diamond 1 981 , 1 983a,b;
Glasser and Marr 1 985 ; Canham 1 987; Canham et al. 1 987;
Kollek et al. 1 986; Duchesne and Bérubé 1 992, 1 994;
Kawamura and Takemoto 1 988; Page and Vennesland 1 983 ;
Kawamura et al. 1 987; Andersson et al. 1 989; Durand et al.
1 990; Yilmaz and Glasser 1 990; Rasheeduzzafar and Hussain
1 991 ; Shayan et al. 1 993 ; Wiens et al. 1 995 ; Nagataki and Wu
1 995 ; Shehata et al. 1 999; Ramlochan et al. 2000; Shehata
and Thomas 2002; Bleszynski 2002; Boddy et al. 2003 ).
Studies on the e ffect o f f y ash and slag on the pore solution
of pastes have been reviewed by Thomas (1 996) and studies
involving silica fume have been reviewed by Thomas and
Bleszynski (2001 ). These studies show that the incorporation of most SCMs leads to a reduction in the concentration
Fig. 5.4.3a‒–Evolution o f the pore solution in pastes
containing SCM (Shehata et al. 1999; Ramlochan et al.
2000; Bleszynski 2002; Shehata and Thomas 2002).
Standard Sharing Group and our chats
Fig. 5.4.3b–‒E ffect o f SCM type and replacement level on
the pore solution hydroxyl ion concentration at 2 years
(1: Shehata and Thomas 2002; 2: Shehata et al. 1999; 3:
Bleszynski 2002; 4: Ramlochan et al. 2000).
of alkali hydroxides in the pore solution of pastes, mortar,
and concretes, with the amount of reduction increasing with
higher SCM contents. Figure 5.4.3a shows the evolution of
the hydroxyl ion concentration of the pore solution extracted
from sealed paste samples with w/cm = 0.50, and Fig. 5.4.3b
shows the OH – concentration at 2 years as a function of the
SCM content (from Thomas [2011 ] , using data from Shehata
et al. [1 999] , Ramlochan et al. [2000] , Bleszynski [2002] ,
and Shehata and Thomas [2002] ). The most e ffcient SCM in
terms of reducing the pore solution alkalinity is silica fume,
followed closely by metkaolin and low-calcium f y ash. Slag
and high-calcium f y ash are less e ffective and have to be
used at higher cement replacement levels. Thomas (201 1 )
showed that the concentration of alkali in the pore solution is
a function of the composition of the binder (cement + SCM),
especially its alkali (Na2 Oe), calcium (CaO), and silica
contents. Figure 5.4.3c shows a strong correlation between
the hydroxyl ion concentration in the pore solution and the
chemical parameter (Na2 Oe· CaO)/(SiO 2 ) 2 calculated from
the chemical composition of the binder (cement + SCM).
Thus, SCMs with higher silica contents and lower calcium
and alkali contents will be more e ffcacious in terms o f
controlling the alkali available for reaction.
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
The ability of SCMs to reduce pore solution alkalinity is
linked to their effect on the composition and alkali-binding
capacity of the hydrates. The introduction of SCMs reduces
the Ca/Si ratio of the hydrates, which results in more alkali
being bound (Bhatty and Greening 1 978 ; Rayment 1 982;
Uchikawa et al. 1 989; Thomas et al. 1 991 ). Glasser and Marr
(1 985) explain the differences in alkali absorption on the
basis of the surface charge on the C-S-H, which is dependent
on the Ca/Si ratio. At high Ca/Si ratios, the charge is positive and the C-S-H tends to repel cations. As the Ca/Si ratio
decreases, the positive charge decreases, becoming negative at Ca/Si ratios less than 1 .3 (Glasser 1 992). Negatively
charged C-S-H has an increased capacity to absorb cations,
especially alkalis. Hong and Glasser (1 999) confrmed the
importance of the Ca/Si ratio on the alkali-binding capacity
of synthesized single-phase C-S-H, but subsequently showed
that the binding capacity could be greatly increased by introducing alumina into the C-S-H to form C-A-S-H (Hong and
Glasser 2002).
SCMs that are highly e ffcient at binding alkalis and
reducing the alkali concentration of the pore solution are
also found to be highly effective in controlling expansion of
concrete containing reactive aggregate. Figure 5.4.3d shows
the 2-year expansion of concrete prisms (ASTM C1 293 ) as
a function of the type and amount of SCM used (Thomas
2011 ). Once again, it was found that the SCMs with the most
silica and the lowest Ca/Si ratio, silica fume, metakaolin,
and low-calcium f y ash are most e ffective and have to be
used at replacement rates of 1 0 to 30 percent, whereas slag
and high-calcium f y ash have to be used at higher replace ment levels. With the exception of materials with very
high alkali contents, all SCMs can be used to control ASR,
provided that they are used at an adequate level of replacement. The amount of SCM required to control ASR depends
on (Thomas 2011 ):
a) Composition of SCM: Increasing amounts are required
as the alkali or calcium content of the SCM increase or as the
silica content decreases.
b) Alkali contributed by the portland cement: Generally increased amounts of SCM are required as the alkali
provided by the cement increases.
c) Reactivity of aggregate: The amount of SCM required
increases as the reactivity of the aggregate increases.
In most conditions, the following levels of replacement
are usually suffcient to control expansion due to ASR:
a) Silica fume: 1 0 to 1 5 percent
b) Metakaolin: 1 5 to 20 percent
c) Low-CaO f y ash: 20 to 30 percent
d) Slag: 35 to 50 percent
e) High-CaO f y ash: greater than or equal to 40 percent
As discussed previously, however, the amount of SCM
required should be determined on a case-by-case basis by
appropriate performance testing, or by reference to prescriptive guidelines developed from empirical data. AASHTO
PP065 and ASTM C1 778 provide both performance-based
and prescription-based methodologies for determining the
required SCM content.
27
Fig. 5.4.3c–‒Relationship between pore solution composi tion and the chemical composition o f the binder.
Fig. 5.4.3d–‒E ffect o fSCMs on 2-year expansion o fconcrete
containing siliceous limestone (1: Shehata et al. 1999; 2:
Shehata and Thomas 2002; 3: Bleszynski et al 2002; 4:
Ramlochan et al. 2000; 5: Thomas and Innis 1998).
5. 4. 4 Use o f chemical admixtures —Chemical admixtures
to inhibit deleterious ASR have not been widely employed
by the construction industry. These include lithium salts,
barium salts, sodium silica f uoride, and alkyl alkoxy silane.
Interest in the use o f lithium compounds, specifcally lithium
nitrate, has resulted in signifcant research and testing in both
laboratory and feld environments. These studies have inves tigated the ability for lithium compounds to control ASR in
new concrete construction as well as the potential of lithium
to reduce ongoing ASR in existing ASR-affected concrete
elements. Examples are pavements, highway barriers, and
bridge elements. A brie f discussion o f the fndings regarding
lithium salts follows.
5. 4. 4. 1
Lithium salts —Although McCoy and Caldwell
(1 951 ) reported on the ability of lithium compounds (LiF,
LiCl, and Li 2 CO 3 ) to control ASR, the use of lithium has not
been adopted by the construction industry, probably due to
its relatively high cost. Interest in the use of lithium has been
shown and a major research project, including feld trials,
was conducted by Stark et al. (1 993).
Initial work by McCoy and Caldwell (1 951 ) and Lawrence
and Vivian (1 961 ) indicated that a level of Li/(Na + K) molar
ratio of 0.74 was necessary to control ASR. Tremblay et al.
(2007) and Feng et al. (2008), however, have shown that
28
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
the dosage of lithium required to control ASR varies greatly
and is largely dependent on the aggregate type. Several
researchers have shown that molar ratios of (Li)/(Na + K) in
the range o f 0.60 to 1 .1 are suffcient to suppress expansion
for many aggregate types (Sakaguchi et al. 1 989; Stark et
al. 1 993 ; Tremblay et al. 2007). Tremblay et al. (2007) have
also shown that for certain aggregate types, such as granitic
gneisses and greywackes, doses as high as (Li)/(Na + K) =
1 .1 might not be suffcient. Additional caution is advised
because insu ffcient lithium may actually increase expansion
(Stark et al. 1 993). Other research has shown that lithium
salts such as LiOH and LiCO 3 are less effective than LiNO 3
in reducing or eliminating ASR (Collins et al. 2004). Several
documents provide more detailed review material and guidance for the use of lithium-based admixtures to control ASR
(AASHTO 2000; Folliard et al. 2006).
5. 4. 4. 2
Other chemical admixtures —Other chemical
compounds were found to reduce expansion due to ASR;
these include various barium salts (Hansen 1 960), sodium
silico f uoride, and alkyl alkoxy silane ( Ohama et al. 1 989).
A wide range of compounds was studied by Hudec and Larbi
(1 989), but the results were largely inconclusive. Before
such admixtures are recommended for commercial applications, further research is required to confrm their e ffcacy
and to elucidate their role in controlling ASR.
5.5—Tests for evaluating preventive measures
5. 5. 1 ASTM C441/C441M—The pyrex mortar bar test, or
morecommonly
FREE standards
from
ASTM C441 /C441 M, Get
has been
used for evaluating the e ffcacy o f pozzolans and slag in controlling expan sion due to ASR. This test method was developed in the
1 940s as a method for assessing the suitability of pozzolans
for use in concrete containing reactive aggregate, such as
in connection with specifcations for siliceous admixtures
for the Davis Dam (Gilliland and Moran 1 949). Early tests
(Buck et al. 1 953 ; Pepper 1 964; Blanks 1 950) indicated that
f y ash and slag were less e ffective than highly siliceous
natural pozzolans, and that they should be used in proportions exceeding 40 percent to be e ffective as defned by
ASTM C441 /C441 M . Since then, numerous other workers
have used this test to evaluate the performance of pozzolans
and slag.
In ASTM C441 /C441 M, the 1 4-day expansion of mortar
bars made with high-alkali cement (0.95 to 1 .05 percent
Na2 O e) and 25 percent f y ash by volume, or 50 percent slag
and stored at 1 00°F (38°C), is compared with that of control
bars (cement only), and the percentage reduction due to the
pozzolan or slag is calculated. Alternatively, the materials
and mixture proportions to be used in the actual job may
be used. ASTM C61 8 requires that the expansion of the test
mixture, regardless of alkali content of cement used, be no
greater than the expansion of a low-alkali control. SCMs
meeting this requirement are considered to be as effective
as the low-alkali cement control for mitigating ASR. The
percentage of pozzolan used in practice is assumed to be
equal to or greater than that used in the test mixture, and also
assumed that the alkali content o f the feld cement used will
not exceed that of the test cement by more than 0.05 percent.
Although ASTM C989/C989M , a specifcation for slag,
does not include a requirement relating to ASR, a nonmandatory appendix suggests the use of ASTM C441 /C441 M
with 1 4-day expansion reduced by 75 percent of control or
kept below 0.02 percent when using project materials. Early
versions of the test required slag to be used at replacements
of 20 percent by volume. The criteria used to assess pozzolans or slag in this test have been criticized as too conservative (Klieger and Gebler 1 987; Kennerley 1 988; Kennerley
et al. 1 981 ; Sturrup et al. 1 983 ).
The potential for silica fume to reduce ASR expansion
is apparent i f used at the specifed 25 percent by volume
replacement in ASTM C441 /C441 M, with shrinkage often
being observed after the normal 1 4-day testing period
(Popovic et al. 1 984). Perry and Gillott (1 985) used various
silica fume contents and found that 1 0 percent was effective
in reducing the 1 4-day expansion by more than 75 percent
compared with control (this was the acceptance criterion at
the time). Other workers have confrmed the ability o f 1 0
percent silica fume (Rasheeduzzafar and Hussain 1 991 ;
Hooton 1 993 ) or less (Bérubé and Duchesne 1 992) to meet
this criterion. Perry and Gillott (1 985), however, observed
continued expansion of the silica fume specimens beyond
1 4 days and questioned the reliability of short-term testing
by this method.
Results from tests with borosilicate glass (ASTM C441 /
C441 M ) have shown the e ffects o f f y ash and slag to vary
considerably between studies. The e ffect o f f y ash has been
Standard
Sharing
chatscalcium content,
shown to
dependGroup
on its and
alkaliour
content,
pozzolanicity, and fneness. The only consensus from the
literature is that the e ffectiveness o f f y ash and slag increases
as the level o f replacement increases, and that all f y ashes
and slags can be used to control reaction, provided that they
are used in suffcient quantity. Pepper and Mather (1 959)
reported the effectiveness of a pozzolan or slag was related
to fneness, alkali release, and the amount o f silica dissolved.
Borosilicate glass is extremely sensitive to test conditions
(surface area, alkali content, and temperature) and contains
signifcant quantities o f alkalis that may be released into
the pore solution. Furthermore, borosilicate glass produces
damaging reaction in just a few days. Consequently, determining the role of pozzolans and slag, particularly their
alkali contributions, is complicated by the use of borosilicate
glass. There has been increased concern over the validity of
ASTM C441 /C441 M (Hobbs 1 989) because the results do
not correlate well with data from concrete tests using natural
aggregates (Bérubé and Duchesne 1 992). Generally, the
replacement level required to limit expansion in the borosilicate glass mortar bar test is signifcantly higher than that
required to limit expansion in concrete containing natural
reactive aggregates; few commercial aggregates are as reactive as borosilicate glass.
ASTM C441 /C441 M is not recommended to determine
the e ffcacy o f lithium nitrate or other lithium compounds
to control ASR, as it does not consider the reactivity of the
aggregate, which is vital to determine the correct dosage
of lithium. In addition, excessive leaching during the test
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
severely limits its ability to reliably predict the e ffcacy o f
lithium nitrate or other lithium compounds.
5.5.2 Accelerated mortar bar test (ASTM C1567) —ASTM
C1 567 was adopted in 2004 to assess the ability of SCMs to
control ASR in mortar bars. The test method is essentially
identical to ASTM C1 260 (5.3.3.3) with the exception that
a portion of the cement is replaced by the SCMs under test.
The expansion limit o f 0.1 0 percent at 1 4 days is specifed
in ASTM C1 567, as it has been correlated with feld performance of concrete (Thomas et al. 2006). While several agencies specify different expansion criteria in terms of percent
expansion reached and test measurement age (1 4 versus
28 days), no additional information has yet been published
to show a stronger correlation to the more reliable ASTM
C1 293 or feld experience with similar combinations o f
aggregates and SCMs.
ASTM C1 567 is not suitable for testing the e ffcacy o f
lithium nitrate, as a signifcant amount o f the originally
incorporated LiNO 3 is leached to the aggressive 1 N NaOH
host solution. Modifed versions o f ASTM C1 567 for eval uating lithium have been proposed (Tremblay et al. 201 0;
AASHTO PP065 ; USACE CRD-C 662).
5.5.3 Concrete prism test (ASTM C1293) —ASTM C1 293
can be used to assess the e ffcacy o f SCMs, chemical admix tures, or both. In this test, a portion of the cement is replaced
by the SCM under evaluation. If a chemical admixture such
as lithium nitrate is to be investigated, a range of dosages
may be necessary to determine the dosage to control ASR
(5.4.4.1 ). The remainder of the testing follows ASTM C1 293
as outlined in 5.3.4.3, with the exception that the duration is
extended to 2 years. Measurements are typically taken at 3or 6-month intervals after the frst year o f testing following
the standard recommendation for concrete prism testing.
An expansion limit o f 0.04 percent at 2 years is specifed
by CSA A23.2-27A and is recommended in the appendix
to ASTM C1 293. This expansion limit has shown a strong
correlation to feld structures cast with similar preventive
measures. This method can also be used to effectively determine the required dosage of lithium salts—namely, lithium
nitrate—to mitigate deleterious ASR.
5.6—Protocols for minimizing the risk o f alkaliaggregate reactivity
Numerous protocols have been developed for minimizing
the risk of alkali-aggregate reactivity in concrete. Many
of these essentially take a two-stage approach. First, the
aggregate is evaluated to determine whether it is potentially alkali-silica reactive or alkali-carbonate reactive.
Second, appropriate measures are selected if the aggregate
is alkali-silica reactive; the aggregate is typically rejected
for use in concrete if it is determined to be alkali-carbonate
reactive. In Canada, both prescriptive (CSA A23.2-27A)
and performance (CSA A23.2-28A) approaches are available for selecting preventive measures. In 201 0, the Canadian protocol was modifed and adopted by AASHTO as
AASHTO PP065. In 201 4, AASHTO PP065 was modifed
and adopted as ASTM C1 778 .
29
The development of AASHTO PP065 is described by
Thomas et al. (2008a) . A fowchart describing the protocol
for determining aggregate reactivity is presented in Fig. 5.6.
Although aggregates can be accepted solely on the basis of
satis factory feld performance petrographic examination, or
both, the protocol warns that a certain level of risk is assumed
by the owner, as either of these approaches may fail to identify a reactive aggregate. Laboratory expansion testing is
recommended and the preferred test is ASTM C1 293, which
is generally believed to be the most reliable for identifying
aggregate reactivity. Because its 1 -year duration renders it
impractical in many situations, however, the protocol does
allow aggregate reactivity to be determined using ASTM
C1 260, recognizing that it frequently results in false positives
in that it identifes some nonreactive aggregates as reactive,
and occasionally results in false negatives in that it fails to
correctly identify some reactive aggregates. Quarried carbonates are evaluated on the basis of their chemical composition
(MgO, CaO, and Al2 O 3 ) to determine the potential for ACR
(CSA A23.2-26A). If the rock is determined to be potentially
alkali-carbonate reactive, it must be tested in concrete, as the
accelerated mortar bar test is not suitable for determining the
risk of ACR. There are three outcomes resulting from the test
protocol shown in Fig. 5.6, and the recommendations for each
outcome are shown in Table 5.6 (CSA A23.2-26A).
Allowable preventive measures include limiting the alkali
content of the concrete, using SCMs, or using lithium-based
admixtures. The level of SCM required can be determined
using either ASTM C1 293 or ASTM C1 567. The dosage
of lithium can also be determined using ASTM C1 293
or a modifed version o f ASTM C1 567. ASTM C1 293
is preferred, but due to its 2-year duration, use of ASTM
C1 567 is also acceptable.
The practice also contains a protocol for determining the
appropriate alkali limit or the level of SCM using a prescriptive approach that has been developed from empirical data.
The limits (maximum alkali content and minimum SCM
level) are based on the criteria:
a) Aggregate reactivity, based on the amount of expansion
in the concrete prism test or accelerated mortar bar test
b) Exposure condition (availability of moisture and
external alkalis) and size of the element
c) Class of structure, based on the required service life and
the consequences should ASR occur
d) Type and composition of SCM
The maximum alkali content of the concrete varies within
the range from 3 to 5 lb/yd 3 (1 .8 to 3.0 kg/m 3 ) Na2 O e,
depending on the risk of ASR and the level of prevention
required. Similarly, the minimum SCM content ranges from
1 5 percent f y ash or 25 percent slag where the risk o f ASR
is low and only mild mitigation measures are required, to
35 percent f y ash or 65 percent slag where there is a high
risk of ASR and more stringent measures are required. In
extreme cases—for example, where a critical structure with
a 1 00-year life is to be built with a highly reactive aggregate
and exposed to alkalis in service—it is necessary to both
limit the alkali content of the concrete and to use high SCM
contents.
30
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
Get more FREE standards from Standard Sharing Group and our chats
Fig. 5. 6‒–Sequence o
f test to
determine aggregate reactivity (CSA A 23. 2-27A ) .
Table 5.6—Testing outcomes and recommended
actions
Outcome of testing
Recommended action
Aggregate is not deleteriously
reactive
Accept aggregate for use – no
prevention required
Aggregate is alkali-carbonate
reactive
Avoid reactive material
Aggregate is alkali-silica reactive
Reject aggregate or select appropriate preventive measures
CHAPTER 6—SULFATE ATTACK
Sulfate attack can take many forms, although it most often
occurs in concrete exposed to external sources of sulfates
(6.1 ). Less commonly, internal sources of sulfate can also
result in damage, particularly when the concrete is exposed to
excessive temperatures at early ages (6.2.2). The attack can be
in the form of chemical attack on the cement paste or physical
attack due to crystallization of sulfate salts (Chapter 8).
6.1 —External sul fate attack
6.1.1
Occurrence —Naturally-occurring
sulfates of
sodium, potassium, calcium, or magnesium that can attack
hardened concrete are sometimes found in soil or dissolved
in groundwater adjacent to concrete structures (Table 6.1 .1 ).
These sulfates have their source from ancient seabed deposits
or a breakdown o f s ulfde or s ul fate- bearing minerals . Indus trial and agricultural e ff uents, as well as municipal was te water, can also supply sulfates. Other sources of sulfates are
water used in concrete cooling towers, where the sulfate ions
gradually
build
up
due
to
evaporation.
S oil
flls
containing
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
industrial waste products, such as slag from iron processing,
can leach sulfate ions.
There are other environments where multiple deterioration mechanisms may be involved. Seawater, brackish
water, and coastal soils constitute a special type of exposure.
Recommendations for these environments are addressed
in 6.3. Portland-cement concrete can also be attacked
by sulfuric acid solutions, which result from oxidation of
sulfur-containing minerals or from decay of organic matter
by bacterial action (Chapter 7).
When water evaporates from a concrete surface, especially
in arid regions, an accumulation of sulfate salts can occur,
resulting in physical salt attack. The general topic of physical salt attack or salt weathering is addressed in Chapter 8.
Sulfate attack has occurred at various locations throughout
the world and is a particular problem in arid areas, such as
the northern Great Plains and parts of the western United
States (Bellport 1 968 ; Harboe 1 982; Reading 1 975 , 1 982;
U.S. Bureau of Reclamation 1 975 ; Verbeck 1 968 ); the
prairie provinces of Canada (Hamilton and Handegord 1 968 ;
Hurst 1 968 ; Price and Peterson 1 968 ); and the Middle East
(French and Poole 1 976). Other non-arid countries, such as
England (Bessey and Lea 1 953 ) and Norway (Bastiensen et
al. 1 957), have also experienced sulfate attack on concrete.
6. 1 . 2
Historical background —Sulfate attack has been
recognized for nearly 250 years. Smeaton (1 791 ), requiring
a material that would harden under water, developed the
frst hydraulic lime formulation o f the industrial age, the
precursor of modern portland cements. While developing a
cement formulation for use in the construction of the Eddystone Lighthouse, Smeaton described its attack by sulfatecontaining solutions in 1 756. During the nineteenth century,
various aspects of sulfate attack of concrete were studied
(Bogue 1 955 ).
Traditionally, sulfate attack has been thought to occur as
a consequence of a sulfate-containing solution entering the
pore structure of concrete and reacting with hydrating cement
compounds such as tricalcium aluminate (C 3 A) to form various
sulfate-containing phases that adversely affect concrete durability. Bates et al. (1 91 3) state that “It is almost universally
believed that it is the reaction of sulphate of magnesia of
the sea water with the lime of the cement and the alumina
of the aluminates of the cement, resulting in the formation
of hydrated magnesia and calcium sulpho-aluminate, which
crystallizes with a large number of molecules of water.”
This understanding has been the basis for the development
of sulfate-resisting cements, and of mixture proportions for
concretes to be placed in sulfate environments.
While there was an understanding that sulfate attack was
associated with a compound frst described by Candlot in
1 880, the precise compositions of calcium sulfoaluminate
hydrates were not established until 1 929. Lerch et al. (1 929)
established the composition of ettringite, known as Candlot’s salt, as 3CaO· Al 2 O 3 · 3CaSO 4 · 31 H 2 O and that of monosulfate as 3CaO· Al 2 O 3 · CaSO 4 · 1 2H 2 O. In addition to establishing the compositions of these salts, Lerch et al. (1 929)
also showed that monosulfate converts to ettringite when
an external source of sulfate is provided. This appears to be
31
Table 6.1.1 ––Mineral names and general
composition o ften used in reports o f sul fate attack
Anhydrite
CaSO 4
Thenardite
Na2 SO 4
Bassanite
CaSO 4 ≈ 0.5 H 2 O
Mirabilite
Na2 SO 4 ≈ 1 0H 2 O
Gypsum
CaSO 4 ≈ 2H 2 O
Arcanite
Kieserite
MgSO 4 ≈ H 2 O
Glauberite
Na2 Ca(SO 4 ) 2
Epsomite
MgSO 4 ≈ 7H 2 O
Langbeinite
K2 Mg 2 (SO 4 ) 3
K2 SO 4
the frst citation in the literature that describes this chemical
mechanism of sulfate attack.
Although the precise composition of ettringite was not
established until the work of Lerch et al. (1 929), the role of
C 3 A in sulfate attack had been recognized earlier. Concern
about the C 3 A content of cements had stimulated a variety
of studies of cement compositions over a period of 20 years
or more. Bates and Klein (1 91 7), who studied the properties of calcium silicates and C 3 A, reached the conclusion that
it would be impossible to commercially produce a portland
cement containing less than approximately 1 percent alumina
because o f the high fring temperatures required. Work in
Canada by Thorvaldsen, however, identifed the means to
reduce C 3 A contents by changing the proportions of C 3 A and
C 4 AF while avoiding excessively high kiln temperatures.
Thorvaldsen observed that cements with high iron contents
also exhibited improved sulfate resistance. This led to the
eventual development of Type V cement (Fleming 1 933 ).
Today, low-C 3 A cements are routinely produced. Sulfate
attack, however, involves phenomena in addition to the
formation of ettringite. Consequently, even so-called zeroC 3 A cements might not be immune to sulfate attack.
The basis of the most commonly used method for establishing the sulfate resistance of cements by measurement
of expansion should also be credited to Thorvaldson et al.
(1 927, 1 929). From observations of deterioration due to
warping and expansion of cement-containing materials, such
as tiles, mortar and concrete, they developed an expansion
test, which is the basis of ASTM C1 01 2/C1 01 2M .
Resistance to sulfate attack is increased by controlling
both cement composition and concrete permeability. The
importance of this was demonstrated in studies by Verbeck
(1 968) and Stark (1 989b) that showed that reduction of
permeability was of greater importance in limiting sulfate
attack than was using a sulfate-resistant cement composition. In a 40-year summary of U.S. Bureau of Reclamation data, it was found that a w/cm of 0.45 or lower helps
in avoiding damage from sodium sulfate attack on portland
cements having C 3 A contents less than 8 percent (Monteiro
and Kurtis 2003 ). In some cases, failure was avoided with a
w /cm as high as 0.53, but signifcant damage can occur in the
w /cm range from 0.45 to 0.53.
An appreciation that limitations on C 3 A content and
on w /cm are both needed to produce sulfate-resistant
concrete have been embodied in various codes and standards governing the selection of concrete for use in sulfate
environments. The U.S. Bureau of Reclamation (1 975) has
formally recognized these requirements since 1 949.
32
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
6.1.3 Mechanisms— Phenomena characterized as external
sulfate attack can occur in a number of ways involving the
formation of a variety of compounds. Origins of sulfates can
be both external and internal, and include the oxidation of
pyrites and generation by bacterial action.
These forms of external sulfate attack have been recognized in Brown (2002):
a) Ettringite (AFt), monosulfate (AFm), and gypsum
formation
b) Sulfate-containing salt formation at or near an evaporative surface
c) Thaumasite formation
Mechanisms for these three forms of external sulfate attack
are described as follows, and publications discussing them
in detail include Lea (1 971 ), Mehta (1 976, 1 992), Mehta
and Monteiro (2006), DePuy (1 994), Taylor (1 997), Hewlett
(1 998), and Skalny et al. (1 998). Publications with particular
emphasis on permeability and the ability of concrete to resist
ingress and movement of water include Reinhardt (1 997),
Hearn et al. (2006), and Diamond (1 998).
6.1.3.1 Sulfate attack associated with ettringite and
gypsum formation —These forms of sulfate attack occur
when
external
sul fate- containing
s olutions
infltrate
the
pores of concrete. This increases the concentration of
sulfates in the concrete pore solution available to react with
sources of calcium and alumina to form ettringite and with
sources of calcium to form gypsum. These reactions are also
in
f
uenced
by
pH
becaus e
ettringite
is
not
s table
below
a
more
FREE
standards
from
pH of approximately 1Get
0 and
gypsum
is not
stable above
a
pH of approximately 1 0.5 (Taylor 1 997). This explains the
common observations
in both laboratory s tudies and in feld
concrete of gypsum near exposed concrete surfaces, where
carbonation has lowered the pH, and of ettringite in the interior of gypsum deposits.
In mature concrete, ettringite typically forms directly
from the monosulfate that had formed during the hydration of the cement. Some supplementary cementitious
materials (SCMs) serve as additional sources of reactive
alumina and thus may increase the potential for ettringite
formation. Low- w/cm concrete mixtures containing SCMs,
however, are generally more resistant to sulfate attack due to
a reduced rate of ingress of the sulfate solution and reduced
CH contents.
Ettringite is responsible for internal cracking and expansion; this damage mechanism is described in early studies
of sulfate attack. However, ettringite may also form as
s econdary
depos its
in
voids
and
cracks
in
mature
feld
concrete exposed to wetting and drying. Consequently, the
presence of ettringite in concrete does not necessarily indicate sulfate attack (ASTM C856).
External sulfate attack produces two damage mechanisms
in concrete. Cracking due to expansion is probably the most
widely reported form of damage. Expansion occurs because
the volumes of ettringite and gypsum are greater than those
of the reactants from which they form. An increase in the
volume of solid phase in the hardened cementitious matrix
results in tensile stresses due to crystallization pressures,
and cracks develop once the tensile strength of the paste is
locally exceeded. A second damage mechanism associated
with external sulfate attack involves softening and loss of
cohesion. As discussed in the following, this damage mechanism involves chemical alterations that destabilize the C-S-H
and calcium hydroxide (CH), and can result in the formation
o f microcracks without signifcant expans ion.
The damage that results from external sulfate attack also
depends on the cation associated with sulfate. The most
common naturally-occurring sulfates that attack concrete
are calcium, sodium, and magnesium sulfate, which are
listed in order of increasing aggressiveness. Calcium sulfate
(gypsum, CaSO 4 · 2H 2 O) is generally the least aggressive
becaus e
its
s olubility
is
s ignifcantly
lower
than
that
of
sodium and magnesium sulfate. Calcium sulfate solutions
can, however, attack concrete (Thorvaldson 1 954; Taylor
1 997; Drimalis 2007). In addition, after calcium and sulfate
ions enter concrete pores, the high alkalinity of the pore
solution increases their solubility compared to that in natural
waters (Hansen and Pressler 1 947), thus allowing the development of higher concentrations of sulfate that can increase
the severity of the attack. The calcium sulfate attack damage
mechanism involves internal expansion and cracking due to
ettringite formation.
A sodium sulfate solution can be more aggressive than
calcium sulfate because sodium sulfate is more soluble.
Consequently, concrete can be exposed to higher sulfate
concentrations. Sodium sulfate attack can lead to the formation of gypsum and ettringite within the cement paste at the
Standard
Group
and our
chatsproducts of CH,
expenseSharing
of the normal
cement
hydration
monosulfate, hydrated C 3 As, and in severe cases, the C-S-H
binder. Both cracking and softening are associated with
sodium sulfate. In addition, salt deposits can form on evaporative surfaces of concrete elements subjected to sodium
sulfate attack, causing scaling (physical salt attack). Chapter
8 discusses the mechanisms of salt crystallization and its
associated damage mechanisms in more detail; understand,
though, that the presence of salt deposits on scaled evaporative surfaces may indicate the occurrence of external chemical sulfate attack, physical salt attack, or both.
Although magnesium sulfate and sodium sulfate share
similar solubility, magnesium sulfate attack can be more
damaging because both magnesium and sulfate ions participate in the attack. The reaction products of magnesium sulfate
attack include ettringite, gypsum, magnesium hydroxide, and
a silica gel, and may produce a matrix with very low strength
or binding capacity (Gollop and Taylor 1 995; Taylor 1 997).
The magnesium ion undergoes a base exchange process
with CH or with C-S-H that forms brucite (magnesium
hydroxide). This lowers the pH of the concrete pore solution and provides a source of calcium to react with sulfate
and produce gypsum. The reactions will continue until they
exhaust the CH and the C-S-H from the paste. Consequently,
softening and loss of cohesion is the end-state damage mechanism associated with magnesium sulfate attack.
6.1.3.2 Physical salt attack by sulfate salts —Deterioration
due to physical salt attack starts at the surface of concrete.
Initially, the deterioration, which has an appearance similar
to scaling caused by freezing and thawing, can be induced
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
by any of several different salts, most commonly sodium
sulfate. Damage is due to crystallization pressure from
the precipitation of crystals within the pore structure of
concrete. Physical salt attack is considered a physical form
of attack because damage is not related to chemical interaction between sulfate ions and the hydrated phases of portland cement. Physical salt attack can occur together with
chemical sulfate attack.
I f sul fate ions are identifed in the soil, groundwater, or as
part o f the e ff orescence on concrete surfaces, any damage o f
concrete by physical salt attack, including damage to architectural surfaces, requires serious consideration and evaluation. Chapter 8 discusses physical salt attack.
6.1.3.3 Thaumasite formation —Although it has a different
chemical composition, thaumasite has a similar crystal structure to ettringite. While it is typical to denote ettringite by
its oxide composition as 3CaO· Al 2 O 3 · 3CaSO 4 · 32H 2 O, the
appropriate representation of the molecular structu re of ettringite is (Ca3 Al(OH) 6 .1 2H 2 O 3+) 2 (SO 4 2–) 3 · 2H 2 O· Ca 3 Al(OH) 6 ;
units form four heavily hydrated columns per unit cell.
There are also four interstices between the columns per unit
cell. Three of these are occupied by sulfate anions and the
fourth by two molecules of water. This structure permits a
broad range of substitutions: divalent cations can be substituted for calcium; trivalent cations, such as transition metal
ions, can substitute for aluminum; tetravalent ions, such
as Si 4+, can also substitute for Al 3+. One such silica-based
compound is thaumasite (CaO.SiO 2 · CaSO 4 · CaCO 3 · 1 5H 2 O
or CaSiO 3 · CaCO 3 · CaSO 4 · 1 5H 2 O).
Thaumasite was frst identifed as occurring in deterio rating concrete in 1 965 by Erlin and Stark (1 966), and later
by Bickley et al. (1 994), and has been extensively investigated by Matthews (1 994), the Thaumasite Expert Group
(1 999) in the UK, and Crammond (2002a,b). Ettringite and
thaumasite are frequently found together in deteriorating
concrete. Whether they are present as intimate mixtures
or exhibit solid solution behavior when formed under
these conditions has been debated. Erlin and Stark (1 966),
however, found lath-shaped crystals where there were petal
overgrowths o f thaumasite on ettringite, fnely banded crys tals of alternating ettringite and thaumasite, and continuous
lath-shaped crystals of which one-half was ettringite and the
other half thaumasite.
Although thaumasite can form in the absence of external
sulfate due to carbonation, which can decompose both ettringite
(releasing sulfates) and the calcium silicate hydrate (C-S-H)
binder (releasing hydrous silica), it is unlikely. External sulfate,
or an inadvertent, gross excess of internal sulfate can also
contribute to thaumasite formation and are almost exclusively associated with its formation. Similarly, if a source of
readily soluble CaCO 3 is present within the concrete, thaumasite formation can occur. Thus, thaumasite can form as a
consequence of sulfate ingress, carbonation, or both. While
originally thought to occur only in concretes exposed to cool
temperatures, thaumasite formation has also been observed
in concrete in temperate climates (Crammond 2002a).
Thaumasite preferentially forms under the cold, wet, alkaline conditions typically experienced by buried concrete
33
structures. The occurrence of thaumasite in deteriorated
building materials has been identifed in a number o f coun tries worldwide, including the United Kingdom, United
States, Canada, South Africa, France, Germany, Norway,
Denmark, Switzerland, Italy, and Slovenia. Probably the
most severe case of thaumasite-damaged concrete encountered so far was in the Canadian Arctic (Bickley et al. 1 994).
The use of sulfate-resisting concrete does not necessarily
prevent the formation of thaumasite, because it is the C-S-H
and not aluminate phases that are attacked by external
sulfates. The replacement of C-S-H by thaumasite transforms the cement paste matrix into a white, soft, noncohesive mass. From the formula for thaumasite, it is seen that
carbonate ions are also necessary.
A positive identifcation o f thaumasite in a cement-based
building material does not automatically indicate that a
problem has occurred or, if it has, that thaumasite was the
cause. There are two distinct ways in which thaumasite
can precipitate as a reaction product within concretes and
mortars (Thaumasite Expert Group 1 999) and the following
characteristics should be considered during diagnosis.
6.1.3.3.1 Thaumasite form of sulfate attack is visually
very distinctive, characterized by signifcant damage to the
cement paste matrix of the concrete or mortar. The main hallmark of thaumasite sulfate attack is that hardened cement
paste becomes partially or totally replaced by thaumasite. As
thaumasite does not possess any binding ability, the affected
cement paste is ultimately transformed into a noncohesive
mass loosely holding the aggregate particles together. Other
distinguishing features include subparallel cracks flled with
thaumasite and white haloes of thaumasite occurring around
aggregate particles. Thaumasite sulfate attack, which causes
gradual softening of the matrix of a buried concrete starting
from the concrete-ground interface and progressing inward,
can sometimes be accompanied by expansive disruption.
6.1.3.3.2 Thaumasite, like ettringite, can precipitate harmlessly in voids and cracks. This phenomenon has been termed
“thaumasite formation” and can be found in concretes or
mortars showing no obvious visual signs of sulfate attack.
Thaumasite formation also occurs in concretes already
damaged by other deterioration mechanisms such as ASR
(French 1 986; Regourd and Hornain 1 986). Although the
presence of thaumasite is more often innocuous, it can be a
precursor to thaumasite sulfate attack.
Thaumasite sulfate attack is typically associated with
ingress of external sulfates. The carbonate ions necessary for
thaumasite sulfate attack can be supplied by limestone aggregates or limestone in cement, or externally by carbonate or
bicarbonate ions dissolved in sulfate-bearing water. Hooton
and Thomas (2002) considered 5 percent limestone additions
to cement not to be a risk. Thaumasite preferentially forms
at temperatures below 59°F (1 5°C). Although it can form
at temperatures up to 77°F (25°C), the rate is much slower
(Thaumasite Expert Group 1 999; Alksnis and Alksne 1 986).
6.1.4 Recommendations
6.1.4.1 Sulfate attack associated with ettringite, mono sulfate, and gypsum formation —Protection against the
various forms of sulfate attack is obtained by proportioning
34
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
Table 6.1.4.1 a—Severity o f exposure conditions
determined from sul fates in soil or water
Exposure class
S0 (not applicable)
Water-soluble sulfate
Sulfate (SO 4 2–) * in
(SO 4 2–) * in soil, %
water, ppm
SO 4
2–
< 0.1 0
SO 4 2– < 1 50
S1 (moderate)
0.1 0 ≤ SO 4 2– < 0.20
1 50 ≤ SO 4 2– < 1 500,
or seawater
S2 (severe)
0.20 ≤ SO 4 2– ≤ 2.00
1 500 ≤ SO 4 2– ≤ 1 0,000
S3 (very severe)
SO 4 2– > 2.00
SO 4 2– > 1 0,000
*
Sulfate expressed as SO 4 is related to sulfate expressed as SO 3 , as given in reports of
chemical analysis of portland cements as follows: SO 3 × 1 .2 = SO 4.
concrete mixtures to minimize the ingress and movement of water using appropriate ingredients. The sulfate
resistance of portland cement generally decreases with an
increase in its calculated tricalcium-aluminate (C 3 A) content
(Mather 1 968 ; Stark 2002). Accordingly, ASTM C1 50/
C1 50M includes Type V sulfate-resisting cement for which
a maximum of 5 percent calculated C 3 A is permitted, and
Type II moderately sulfate-resisting cement for which the
calculated C 3 A is limited to a maximum of 8 percent. There
is also evidence that the alumina in the aluminoferrite phase
of portland cement can participate in sulfate attack (Brown
et al. 1 986). Therefore, ASTM C1 50/C1 50M provides that in
Type V cement, the C 4 AF + 2C 3 A content should not exceed
25 percent unless the alternate requirement based on the use
of the performance test (ASTM C452/C452M ) is invoked.
In the case of Type V cement, the optional sulfate-expanGet morecan
FREE
standards
sion test (ASTM C452/C452M)
be used
in place offrom
the
chemical requirements (Mather 1 978 ). In CAN/CSA A3000,
ASTM C452/C452M expansion limits are used to qualify
both moderate- and high-sulfate-resisting portland cements.
The use of ASTM C1 01 2/C1 01 2M is discussed by Patzias
(1 991 ). ASTM C11 57/C1 1 57M , ASTM C595/C595M , and
CAN/CSA A3000 blended cements also use ASTM C1 01 2/
C1 01 2M expansion limits to qualify moderate (MS) or
high (HS) resistant performance. Both ACI 31 8 and CSA
A23.1 -1 4/CSA A23.2 allow the use of other combinations
of cementing materials in sulfate exposure, provided that
performance testing using ASTM C1 01 2/C1 01 2M demonstrates that the expansion limit for the appropriate exposure
class (Table 6.1 .4.1 a) is not exceeded. Note that in Table
6.1 .4.1 b, performance testing is not required if ASTM C1 50/
C1 50M Type II cement is selected for S1 exposure and
Type V cement is selected for S2 exposure.
One strategy for reducing the ingress and movement of
dissolved sulfates and water is to lower the w/cm . The use
of acceptable SCMs is another, complementary strategy to
reduce the ingress and movement of water into the concrete
(refer to Chapter 3 regarding limiting f uid ingress). Care
should be taken to ensure that the concrete is designed and
constructed to minimize shrinkage cracking. Proper placement, compaction, fnishing, and curing o f concrete are
essential to minimize the ingress and movement o f f uids
that carry aggressive salts. Recommended procedures for
these are found in ACI 304R, ACI 302.1 R, ACI 308R, ACI
305R, ACI 306R, and in Chapter 3 of this guide.
Recommendations for the maximum w/cm and the type
of cementitious material for concrete that will be exposed
to sulfates in soil or groundwater are given in Table 6.1 .4.1 b
for exposures defned in Table 6.1 .4.1 a. Both recommenda tions are important, as limiting only the type of cementitious material is not adequate for satisfactory resistance to
sulfate attack (Kalousek et al. 1 976; Stark 2002; Monteiro
and Kurtis 2003).
Table 6.1 .4.1 b provides recommendations for various
degrees of potential exposure. These recommendations are
designed to protect against distress from sulfate sources
external to the concrete, such as may be in adjacent soil,
groundwater, and e ff uents carried in concrete pipes.
The feld conditions o f concrete exposed to sul fate are
numerous and variable. The aggressiveness of the conditions depends on, among other things, soil saturation, water
movement, ambient temperature and humidity, concentration of sulfate, and type of sulfate or combination of sulfates
involved. Table 6.1 .4.1 b provides criteria that should maximize the service life of concrete subjected to aggressive
sulfate exposure conditions.
6.1.4.2 Physical sulfate attack: physical salt attack by
sulfate salts —Chapter 8 provides recommendations for the
more general case of physical salt attack. In the presence of
sulfates, however, the code requirements for prevention of
chemical sulfate attack must be followed (Tables 6.1 .4.1 a
and 6.1 .4.1 b).
6.1.4.3 Thaumasite formation —Use of Types II or V
Standard
Sharingcements
Groupdoes
andnot
our
chatsthaumasite sulfate
sulfate-resisting
prevent
attack because thaumasite does not consume aluminate
phases. Ettringite, however, does appear to be a precursor
to thaumasite formation in many cases. Thaumasite attack
is sometimes associated with: 1 ) excess carbonate fnes in
aggregates or cements that are well above the concentrations allowed by ASTM C1 50/C1 50M; and 2) carbonation
from air or water exposures. The conditions under which
thaumasite attack occurs are not fully known and, as yet,
there are no standards that address its prevention specif cally. Very wet exposure conditions, however, appear to
be common with thaumasite sulfate attack, so provision of
low- w/cm concretes would reduce the ingress of sulfate and
carbonate ions as well as reducing the rate of carbonation
(Hooton 2007). ASTM C1 01 2/C1 01 2M, when modifed
such that mortar bars are exposed to the sulfate solution at
40°F (5°C), has been found to be suitable for determining
whether cementitious binders are resistant to thaumasite
sulfate attack (Hooton and Brown 2009; Hooton et al. 201 0);
this test was adopted in CSA A3000 in 201 0. In addition to
measuring length and mass change, X-ray diffraction can
determine whether thaumasite is present.
The use of slag cement in concrete appears to help resist
thaumasite sulfate attack (Hill et al. 2003). Nobst and Stark
(2003) found that concrete with cement containing at least
66 percent slag cement was resistant to thaumasite. This was
confrmed by Bellmann and Stark (2008) where mortar bars
made with a European CEM IIIB i cement with 65 percent
slag cement was resistant to thaumasite after exposure to
sulfate solutions at 46°F (8°C). They attributed the good
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
35
Table 6.1.4.1 b—Requirements to protect against damage to concrete by sul fate attack from external
sources o f sul fate
Prescriptive cementitious material requirements
Performance cementitious material requirements
Maximum expansion when tested using ASTM
Severity of
Cement types *
C1012/C1012M
potential
w/cm by mass,
ASTM
ASTM
ASTM
exposure
maximum
C150/C150M
C595/C595M
C1157/C1157M
At 6 months
At 12 months
At 18 months
S0
No w/cm
restriction
No type
restriction
No type restriction
No type
restriction
—
—
—
S1
0.50 †
Type II ‡§
IP (MS), IS (<70) (MS),
IT (P<S<70) (MS), or IT
(P≥S) (MS)
MS
0.1 0%
—
—
S2
0.45 †
Type V #
IP (HS), IS (<70) (HS),
IT (P<S<70) (HS), or IT
(P≥S) (HS)
HS
0.05%
0.1 0% ||
—
S3
0.40 †
Type V plus
pozzolan or
slag cement **
IP (HS), IS (<70) (HS),
IT (P<S<70) (HS), or IT
(P≥S) (HS)
HS ††
—
—
0.1 0%
*
Alternative combinations of cementitious materials to those listed in Table 6.1 .4.1 b can be permitted when tested for sulfate resistance and meeting the ASTM C1 01 2/C1 01 2M
expansion criteria for the severity of potential exposure.
‡
Other available types of cement, such as ASTM C1 50/C1 50M Type I or Type III can be permitted in Exposure Classes S1 if the C 3 A content is less than 8 percent.
§
For seawater exposure, other ASTM C1 50/C1 50M cement types with C 3 A contents up to 1 0 percent are permitted if w/cm does not exceed 0.40. (Refer to Section 6.3 on seawater
exposure.)
#
An ASTM C1 50/C1 50M Type III cement with the optional limit of 5 percent can be permitted or ASTM C1 50/C1 50M cement of any type having expansion at 1 4 days no greater
than 0.040 percent when tested by ASTM C452/C452M.
||
The 1 2-month expansion limit can be used if the 6-month limit is not met, but is not required if the 6-month limit is met.
†
Values applicable to normalweight concrete. They are also applicable to structural lightweight concrete except that the maximum w/cm of 0.50, 0.45, and 0.40 should be replaced
by specifed 28-day compressive strengths o f 26, 29, and 33 MPa (3750, 4250, and 4750 psi), respectively.
**
As stated in ACI 31 8, the amount o f the specifc source o f the pozzolan or slag cement to be used shall be at least the amount that has been determined by service record to
improve sul fate resistance when used in concrete containing Type V cement. Alternatively, the amount o f the specifc source o f the pozzolan or slag cement to be used shall be at
least the amount tested in accordance with ASTM C1 01 2 and meeting the criteria shown in the table.
††
For Exposure Class S3, ASTM C11 57/C11 57M HS cement must contain pozzolan cement, slag cement, or both.
resistance to the reduction in CH content of the matrix, and
when CH was added to mortar bars made with the same
cement, damage due to thaumasite sulfate attack occurred.
Fifty percent slag will resist the formation of thaumasite
in mortar bars stored at 40°F (5°C) (Hooton et al. 201 0).
Bellmann and Stark (2007) also found that when a CEMIIAL cement, which is portland cement with up to 20 percent
interground limestone, was replaced with either 20 or 40
percent Class F f y ash, there was no damage after 4.5 years
of storage in 1 500 mg/L SO 4 = (sodium sulfate) solution at
46°F (8°C).
6.1.5 Sampling and testing to determine potential sulfate
exposure —To assess the severity of the potential exposure
of concrete to detrimental amounts of sulfate, representative samples should be obtained o f both the f uid and sul fate
compound(s) that might reach the concrete or of soil that
might be leached by water moving to the concrete. The
procedure for making a water extract of soil samples for
sulfate analysis that is given in ASTM C1 580 is recommended (Hayes 2007). Although other methods have been
used, the results are often affected by the test method, especially the extraction ratios.
6.1.6 Establishing equivalent performance for cementi tious materials —The use of alternative combinations of
cementitious materials to those listed in Table 6.1 .4.1 b is
permitted for any class of exposure. Any binary or ternary
blend of portland cement of any type meeting ASTM C1 50/
C1 50M, ASTM C595/C595M, or ASTM C11 57/C1 1 57M
with f y ash or natural pozzolan meeting ASTM C61 8, silica
fume meeting ASTM C1 240, or slag cement meeting ASTM
C989/C989M is permitted if it meets the expansion limits
in Table 6.1 .4.1 b when tested in accordance with ASTM
C1 01 2/C1 01 2M.
The portland-cement portion of the test mixture should
always consist of cement with Bogue-calculated C 3 A content
of not less than that being proposed for use. Material qualifcation tests using the expansion limits in Table 6.1 .4.1 b
should be based on passing results from two samples taken
at times a few weeks apart. The qualifying test data should
be no older than 1 year from the date of test completion.
6.1.7 Proportions and uni formity o f pozzolans and slag
cement —The proportion o f f y ash, natural pozzolan, silica
fume, or slag cement used in the project mixture (in relation
to the amount of portland cement) should be the same as
that used in the test mixture prepared to meet the recommendations of Table 6.1 .4.1 b and Section 6.1 .6. In blends
or mixtures with portland cement containing only one SCM,
such as f y ash, natural pozzolan, silica fume, or slag cement,
the proportion o f f y ash or natural pozzolan can generally
be expected to range between 1 5 and 50 percent by mass
of the total cementitious material, depending on the severity
of exposure. Similarly, the proportion of silica fume can be
expected to range between 5 and 1 2 percent by mass of the
total cementitious material, and the proportion of cement
36
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
can be expected to range between 35 and 70 percent by
mass of the total cementitious material. When more than one
supplementary cementitious material is used, the individual
proportion of each may be less than these values.
The uni formity o f the f y ash or s lag cement us ed in the
project should be within the following limits compared to
that used in the mixtures tested to meet the recommendations of Table 6.1 .4.1 b and Section 6.1 .6:
a) Fly ash: reported calcium-oxide content (analyzed in
accordance with ASTM C1 1 4) no more than 2.0 percentage
points higher than that o f the f y as h us ed in the tes t mixture
b) Slag cement: reported aluminum-oxide content
(analyzed in accordance with ASTM C11 4) no more than 2.0
percentage points higher than that of the slag cement used in
the test mixture.
The portland cement used in the project should have a
Bogue-calculated C 3 A content no higher than that used in
the mixtures tested to meet the recommendations of Table
6.1 .4.1 b and Section 6.1 .6.
Studies have shown that some pozzolans and slag cements
used, either in blended cement or added separately to the
concrete in the mixer, increase the life expectancy of concrete
considerably in sulfate exposure. Many slag cements and
pozzolans signifcantly reduce the permeability o f concrete
(Bakker 1 980; Mehta 1 981 ). They also combine with the
alkalis and CH released during the hydration of the cement
(Vanden Bosch 1 980; Roy and Idorn 1 982; Idorn and Roy
1 986), reducing the potential for gypsum formation (Lea
Get more
standards
1 971 ; Biczok 1 972; Kalousek
et al.FREE
1 972; Mehta
1 976). from
Table 6.1 .4.1 b requires a suitable pozzolan or slag cement
along with Type V cement or equivalent in S3 exposures.
Research indicates that some pozzolans and slag cements
are effective in improving the sulfate resistance of concrete
made with Type I and Type II cement; this option is allowed
if the 1 8-month ASTM C1 01 2/C1 01 2M expansion limit in
Table 6.1 .4.1 b is met. Some pozzolans, especially Class C
f y as hes , decreas e the sul fate res is tance o f mortars in which
they are used (Mather 1 981 b; Mather 1 982). Good results
were obtained when the pozzolan was a
f
y as h meeting the
requirements ofASTM C61 8 Class F (Dikeou 1 975; Dunstan
1 976). Slag cement should meet ASTM C989/C989M and
silica fume should meet ASTM C1 240.
In concrete that is made with non-sulfate-resisting
cements, calcium chloride reduces resistance to attack by
sulfate (U.S. Bureau of Reclamation 1 975) and, therefore, its
use should be prohibited in concrete exposed to sulfate (S-1
or greater exposure). If Type V cement is used, however, it
is not harmful to use calcium chloride in normally acceptable amounts as an accelerating admixture (Mather 1 992).
Calcium chloride, however, can induce and accelerate corrosion of reinforcing steel and aluminum conduit.
6.2—Internal sul fate attack
materials
sul fate
contents that are deleterious to concrete. Allowable sulfate
contents in cements meeting ASTM C1 50/C1 50M were
increased several times from 1 941 to 1 971 (Hooton 2008), as
6. 2 . 1
meeting
Concrete
materials
current AS TM
—Cementitious
s pecifcations
will
not
have
cement compos itions
and fnenes ses changed to allow better
optimization of sulfate contents. ASTM C1 50/C1 50M now
allows SO 3 limits to be exceeded if it can be demonstrated
(typically using ASTM C563) that the optimum SO 3 content
is above the stated limit. In this case, ASTM C1 038/C1 038M
must show that the SO 3 content of the cement will not result
in adverse expansions. These results are considered satisfactory when a 1 4-day expansion limit of 0.020 percent is specifed.
This tes t and expans ion limit has als o been adopted in
ASTM C1 1 57/C11 57M and is used for Canadian portland
cements, blended cements, and combinations of cementitious materials in CSA A3001 (Hooton and Brown 2009).
This test and expansion limit has not been adopted in ASTM
C595/C595M because it still relies on ASTM C265.
Although ASTM C33/C33M does not limit sulfate content
in aggregates for use in concrete, they should not contain
appreciable levels of sulfate-bearing minerals, such as
calcium sulfate inclusions, or be contaminated with sulfates
or
s ulfdes
air- cooled,
s lag
s uch
as
blas t-
aggregate
pyrite.
furnace
include
Limits
s lag
iron
are
placed
aggregate.
s ulfde
and
on
The
s ulfdes
s ulfdes
calcium
in
in
s ulfde.
Aggregate sampling and testing for sulfate content should
be completed in advance of use. Sulfate concentrations in
mixing water are not normally deleterious to concrete and
should meet the limits in ASTM C1 602/C1 602M. Sulfates in
chemical admixtures meeting ASTM C260/C260M, ASTM
C494/C494M, and ASTM C1 01 7/C1 01 7M will not be deleterious to concrete.
Standard Sharing
Group and
our chats
Delayed ettringite
formation
(DEF)
6. 2 . 2
6. 2 . 2 . 1
Occurrence —Under
certain conditions, heatcured concrete elements can suffer expansion and cracking
on subsequent exposure to moisture. This form of deterioration has been commonly referred to as delayed-ettringite
formation (DEF). The normal early formation of ettringite
that occurs in concrete cured at ambient temperature can be
delayed as a result of exposure to excessive temperatures
during manufacture. The ettringite then forms at later ages
when the concrete is exposed to moisture in service. This
delayed formation of ettringite can lead to internal expansion
and damage in hardened concrete. In the 1 990s, it appeared
that heat-cured railway ties were particularly susceptible to
this form of deterioration with cases involving ties being
reported in Germany, Finland, former Czechoslovakia,
Canada, the United States, South Africa, and Australia (Heinz
and Ludwig 1 987; Tepponen and Eriksson 1 987; Vitousova
1 991 ; Mielenz et al. 1 995; Oberholster et al. 1 992; Shayan
and Quick 1 992). In most, if not all, of the alleged cases of
DEF, other mechanisms of deterioration, especially alkalisilica reaction (ASR) and freezing-and-thawing damage,
have als o been implicated, making it di ffcult to identi fy the
precise role of DEF in the deterioration.
The occurrence of these problems and apparent role of
elevated-temperature curing has led many countries to
impose restrictions on heat curing of precast concrete. These
restrictions include limits placed on preset times, rates of
heating and cooling, and maximum temperature. There is
evidence that these and similar practices in Europe have
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
been successful in eliminating damage due to DEF (Skalny
and Locher 1 997; Stark and Bollmann 1 999).
Note that the risk of damage due to DEF is not restricted
to heat-cured precast concrete elements. Internal concrete
temperatures may increase suffciently, due to the heat
released during the hydration of the cementitious component
of the concrete, to promote DEF in non-heat-cured precast
concrete or in massive cast-in-place concrete elements
(Thomas et al. 2008b). Indeed, the new European practices
impose similar limits on the maximum internal concrete
temperature for both precast and cast-in-place concrete.
6.2.2.2 Mechanisms —As reviewed by Thomas and Skalny
(2006), the solubility of ettringite increases with temperature
and pH, and elevated temperatures increase both the sulfate
and alumina concentrations in the pore solution of concrete.
Much of this sulfate and alumina is taken in by the rapidly
forming calcium-silicate hydrates (C-S-H). Immediately
after the early exposure to elevated temperature, little or no
ettringite is detected in the concrete, and poorly crystalline
monosulfate appears to be the main sulfate-bearing phase
(Taylor et al. 2001 ). During subsequent exposure to moisture at ambient temperatures, most of the sulfate, but only
a small amount of the alumina, is released by the C-S-H.
The increased availability of sulfate results in a conversion
of the monosulfate into ettringite, which could occur many
months or years following the exposure to elevated temperatures. This delayed formation of ettringite may, under some
circumstances, result in expansion of the paste and consequent cracking of concrete. The cement paste expansion is
generally believed to be a result of the growth of ettringite
crystals in the very small pores (approximately 1 00 nm) of
the cement paste (Taylor et al. 2001 ; Lawrence 1 995). The
expansion of cement paste results in the formation of gaps
around aggregate particles and cracking of the cement paste
(Johansen et al. 1 993). Ettringite eventually reprecipitates
into these gaps and the cracks, but this is not a cause of
damage (Johansen et al. 1 993).
Reviews of laboratory studies (Day 1 992; Thomas 2001 )
and recent data (Thomas et al. 2008b) indicate that expansion due to DEF is unlikely to occur unless mortar or
concrete specimens are subjected to elevated temperatures
in excess of approximately 1 60°F (70°C) and that the risk of
expansion increases with increasing temperature above this
threshold value. For mortar or concrete that has been exposed
to higher temperatures, the risk of expansion appears to be
a function of many parameters, including both physical
and chemical characteristics of the cementitious binder
(Shimada 2005; Shimada et al. 2007). A review of published
results (Thomas 2001 ) appears to indicate that high-fneness
cements produced from clinker with high concentrations of
C 3 A, C 3 S, and Na2 O eq, and consequently having a high SO 3
content, have the greatest susceptibility to DEF expansion
when heat cured. Kelham (1 996) showed a clear correlation
between the 2-day compressive strength of mortar and the
expansion of the same mortar subsequent to heat curing at
1 94°F (90°C). Although studies have shown relationships
between DEF expansion and various compositional parameters of the cement, there is no single parameter that can
37
be used to reliably predict the performance of a particular
cement. Thus, it is not possible to impose a single limit on
the chemical composition of the cement to eliminate the risk
of expansion in concrete that may be exposed to excessive
temperature during curing. However, it is apparent from
laboratory studies that portland cements having high C 3 A,
C 3 S, Na2 O eq, and SO 3 contents generally have the highest
propensity for expansion when cured at high temperatures.
Cements that exhibit lower early-age strength development
generally present a lower risk of expansion (Kelham 1 996;
Ramlochan 2002).
The risk of expansion of heat-cured mortars and concrete
can be effectively eliminated by the incorporation of enough
of the appropriate SCMs (Ghorab et al. 1 980; Ramlochan et
al. 2003). Silica fume, when used as the sole SCM, reduces
but does not fully mitigate DEF, apparently due to the lack of
aluminates in the hydrates (Ramlochan et al. 2003). This is
the reason that the recommendations shown in Table 6.2.2.2
require silica fume to be used in a ternary system with f y ash
or slag cement.
6.2.2.3 Recommendations —To minimize the risk of
poor durability due to deleterious DEF reactions associated with exposure to elevated temperatures at early ages,
the maximum internal temperature of concrete should be
controlled such that it does not exceed 1 58°F (70 o C) at any
time. If temperatures in the range o f 1 58°F < T ≤ 1 85°F (70
to 85°C) are unavoidable, the measures in Table 6.2.2.2
should be adopted.
6.3—Seawater and brine exposure
6.3.1 Occurrence —Seawater throughout the world varies
in the concentration of total salts. The proportions of the
constituents of seawater salts, however, are essentially
constant. More concentrated brines are contained in some
land-locked bodies of water, such as the Great Salt Lake and
the Dead Sea. Brackish water is also aggressive to reinforced
concrete.
The concentration is lower in the colder and temperate
regions than in warm seas and is especially high in shallow
coastal areas with high evaporation rates. The concentration
of salts in land-locked seas also depends on the amount of
fresh water f owing in from rivers (Hewlett 1 998).
Where concrete structures are placed on reclaimed coastal
areas with the foundations below saline groundwater levels,
capillarity and evaporation may cause supersaturation and
crystallization of salts in the concrete above ground, resulting
both in chemical and physical attack on concrete from
sulfates, and in aggravated corrosion of steel from accompanying chlorides. Where concrete structures are immersed in
sea water, the portions above water are usually affected the
most by sulfate attack, both physical and chemical, while
portions that are totally immersed often suffer considerably
less damage (Hewlett 1 998).
These combined deleterious effects can cause severe
defects in concrete in the course of a very few years, especially in tropical climates where high temperature increases
the rate o f deterioration. This section focuses on the inf u-
38
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
Table 6.2.2.2—Recommended measures for reducing potential for DEF in concrete exposed to elevated
temperatures at early ages*
Maximum concrete temperature
Prevention required
T ≤ 1 58°F (70°C)
T ≤ 1 85°F
T ≤ 85°C)
No prevention required
T > 1 85°F (85°C)
The internal concrete temperature should not exceed 1 85°F (85°C) under any circumstances.
1 58°F <
(70°C <
*
T
Use one of the following approaches to minimize the risk of expansion:
1 . Portland cement meeting requirements of ASTM C1 50/C1 50M moderate or high sulfate-resisting and low-alkali
cement with a fneness value less than or equal to 430 m2 /kg
2. Portland cement with a 1 -day mortar strength (ASTM C1 09/C1 09M) less than or equal to 2850 psi (20 MPa)
3. Any ASTM C1 50/C1 50M portland cement in combination with the following proportions of pozzolan or slag
cement:
a) Greater than or equal to 25 percent f y ash meeting the requirements o f ASTM C61 8 for Class F f y ash
b) Greater than or equal to 35 percent f y ash meeting the requirements o f ASTM C61 8 for Class C f y ash
c) Greater than or equal to 35percent slag cement meeting the requirements of ASTM C989/C989M
d) Greater than or equal to 5 percent silica fume (meeting ASTM C1 240) in combination with at least 25 percent
slag cement
e) Greater than or equal to 5 percent silica fume (meeting ASTM C1 240) in combination with at least 20 percent
Class F f y ash
f) Greater than or equal to 1 0 percent metakaolin meeting ASTM C61 8
4. An ASTM C595/C595M or ASTM C11 57/C11 57M blended hydraulic cement with the same pozzolan or slag
cement content as listed in Item 3
Assembled from Ghorab et al. (1 980), Ramlochan et al. (2003), Thomas (2001 ), Thomas et al. (2008b).
ence o f sul fates in seawater. Section 7.2 describes the inf uence of chloride and magnesium ions in seawater.
6.3.2 Mechanisms —The reaction of mature concrete with
sulfate ions in seawater is similar to that with sulfate ions in
fresh water or leached from soils, but the effects are different
(Mather 1 966). Concrete in seawater often exhibits erosion,
softening, or loss in mass as a result of sulfate attack as
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opposed to the expansion, which may also occur in nonsaline sulfate environments.
The presence of chloride ions, however, alters the extent
and nature of the chemical reaction so that less expansion
is produced by a cement of given Bogue-calculated C 3 A
content than would be expected of the same cement in a
freshwater exposure where the water has the same sulfate
ion content. To an extent, this can be explained by the ability
of chlorides to bind with C 3 A in the cement to form chloroaluminates, such as Friedel’s salt (Verbeck 1 975). Formation
of chloroaluminates does not result in undesirable expansion, and it also lowers the amount of C 3 A available to react,
reducing the damage caused by sulfate attack. In the tidal
and splash zones, however, the concentration of sulfate and
chloride ions in concrete can be increased by capillary action
and evaporation.
It has been suggested that the magnesium sulfate in
seawater is primarily responsible for the chemical reactions occurring in concrete. Because CH and calcium sulfate
are both more soluble in seawater than in fresh water, they
are more easily removed by leaching, while magnesium
sulfate forms gypsum, silica gel, and magnesium hydroxide
(Hewlett 1 998). Sulfate attack in seawater can also lead
to decalcifcation o f C-S-H. The formation o f magnesium
hydroxide in concrete pores can act as a barrier to the ingress
of sulfate ions, but this effect is not as pronounced in more
permeable concrete (Hewlett 1 998; Santhanam et al. 2006).
6.3.3 Recommendations relative to seawater or brine —
The rate of deterioration depends on the concentration of
aggressive ions, duration of exposure, and permeability and
chemical resistance of concrete. Low permeability plays an
important role in hindering the ingress of aggressive ions
in seawater or brine. As with conventional sulfate attack,
permeability is more important than chemical composition
of cement in avoiding damage from seawater exposure.
Recommended w/cm and cement types for seawater expoStandard
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sure are provided in Table 6.1 .4.1 b under S1 exposure.
The performance of concretes continuously immersed in
seawater made with ASTM C1 50/C1 50M cements having
C 3 A contents as high as 1 0 percent has proven satisfactory,
provided the permeability of the concrete is low (Browne
1 980). The U.S. Army Corps of Engineers (USACE) (1 984)
and the Portland Cement Association recommend up to
1 0 percent Bogue-calculated C 3 A for concrete that will
be permanently submerged in seawater if the w/cm is kept
below 0.45 by mass.
CSA A23.1 -1 4/CSA A23.2 allows w/cm of up to 0.50 and
recommends 8 percent maximum C 3 A content. With reinforced concrete construction exposed to seawater, however,
the maximum w/cm would be limited to 0.40 due to the
additional chloride exposure. In addition, low C 3 A content
decreases resistance to chloride penetration; therefore, CSA
A23.1 -1 4/CSA A23.2 suggests that C 3 A contents of portland
cement be kept between 4 and 8 percent to help protect the
reinforcement by increasing chloride binding.
Verbeck (1 968) and Regourd et al. (1 980) showed,
however, that there may be a considerable difference
between the calculated and the measured phase composition
of cement, especially as far as C 3 A and C 4 AF are concerned.
Therefore, the interrelation between the measured C 3 A
content and seawater resistance may be equally uncertain.
The requirement for low permeability is essential not
only to delay the effects of sulfate attack but also to afford
adequate protection to reinforcement with the minimum
concrete cover as recommended by ACI 357.1 R for expo-
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
39
Table 7.1 a—E ffect o f common chemicals on concrete *
Rate of attack at
ambient temperature
Inorganic acids
Organic acids
Alkaline solutions
Salt solutions
Miscellaneous
Rapid
Hydrochloric Nitric
Sulfuric
Acetic
Formic
Lactic
—
Aluminum chloride
—
Moderate
Phosphoric
Tannic
Sodium hydroxide
greater than 20 percent
Ammonium nitrate
Ammonium sulfate
Sodium sulfate
Magnesium sulfate
Calcium sulfate
Slow
Carbonic
—
Sodium † hydroxide
1 0 to 20 percent
sodium hypochlorite
Ammonium chloride
Magnesium chloride
Sodium cyanide
Chlorine (gas)
Seawater
Soft water
—
Oxalic
Tartaric
Sodium hydroxide
Less than 1 0 percent
sodium hypochlorite
ammonium hydroxide
Calcium chloride
Sodium chloride
Zinc nitrate
Sodium chromate
Ammonia (liquid)
Negligible
†
*
Refer to Portland Cement Association (2001 ) for a more complete list of chemicals and their potential effects on concrete.
†
The effect of potassium hydroxide is similar to that of sodium hydroxide.
sure to seawater. The required low permeability is attained
by using concrete with a low w/cm and is well consolidated
and adequately cured. ACI 357.1 R recommends a maximum
w/cm of 0.45 for the submerged zone and 0.40 for the splash
zone.
The permeability of concrete made with appropriate
amounts of suitable slag cement or pozzolan can be as low as
1 /1 0 or 1 /1 00 that of comparable concrete of equal strength
made only with portland cement (Bakker 1 980). The satisfactory performance of concretes containing slag cement in
a marine environment has been described (Lea 1 971 ; Vanden
Bosch 1 980; Mather 1 981 a).
Concrete should be designed and constructed to minimize
crack widths, therefore limiting chloride penetration to reinforcement and avoiding the concentration of sulfates. Additionally, concrete should reach a maturity equivalent of not
less than 5000 psi (35 MPa) at 28 days when fully exposed
to seawater.
CHAPTER 7—CHEMICAL ATTACK
7.1 —General
Concrete is rarely attacked by chemicals in their solid
form. To produce a signifcant attack on concrete, aggress ive
chemicals must be in solution and above some minimum
threshold concentration to drive the chemical reactions
that diminish its engineering properties. Although concrete
may perform satisfactorily in a variety of exposure conditions where aggressive chemicals are present, some kinds of
chemical environments
will s ignifcantly s horten the s ervice
fe o f even the best concrete unles s specifc measures are
taken.
An understanding of these exposure conditions permits
measures to be taken to prevent or slow deterioration. Lower
permeability will hinder the infltration o f aggress ive chemi cals. Less-reactive paste can be effective in mitigating deterioration. Concrete members exposed to aggressive solutions
that are under hydraulic pressure from one side may be more
li
Bromine (gas)
S ulfte liquor
vulnerable because the hydraulic gradient can accelerate the
f the aggres s ive s olution into the concrete.
This chapter discusses aggressive chemical exposures,
including: seawater, acids, fresh water, carbonation, industrial chemicals, deicing chemicals, and environmental
structures. Some useful summaries of the potential effects
of chemical exposures include Lea (1 971 ), Biczok (1 967),
Scrivener and Young (1 997), Hewlett (1 998), Eglinton
(1 998), Portland Cement Association (2001 ), ACI 51 5.2R,
and ACI 350.1 . Table 7.1 a summarizes the effects of the
more common chemicals that lead to the deterioration of
concrete. Table 7.1 b summarizes factors that may affect the
rate of chemical attack.
infltration o
7.2—Seawater
7.2.1 Occurrence —Seawater contains dissolved salts that
are potentially aggressive to concrete. The major chemical
components include, in approximate order of decreasing
concentration: chloride, sodium, sulfate, magnesium,
calcium, and potassium. The concentration of total salts
in seawater varies; warmer climates generally have higher
concentrations. The severity of marine exposures can vary
greatly within a given concrete structure. In general, continuous submersion is the least aggressive exposure. Areas
where capillary suction and evaporation are prevalent are
the most aggressive because these processes tend to increase
the concentration of salts. Examples of such exposures
include reclaimed coastal areas with foundations below
saline groundwater level, intertidal zones, and splash zones.
This s ection mainly focus es on the in f uence o f chloride and
magnes ium ions in seawater. S ection 6 . 3 des cribes the in f uence of sulfates in seawater.
7.2.2 Reaction mechanisms —The chemical reactions
that affect concrete exposed to seawater can be complex.
The
mos t
s ignifcant
aggres s ive
s pecies
include
magne sium, sulfate, and chloride. Magnesium ions may react with
calcium hydroxide (CH) and form magnesium hydroxide
or brucite (Lea 1 971 ). This phase is highly insoluble and
40
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
Table 7.1 b—Factors in f uencing chemical attack on concrete
Factors that accelerate or aggravate attack
*
†
‡
#
§
||
Factors that mitigate or delay attack
1 . High permeability due to:
a) High water absorption
b) High w/cm
c) Poor consolidation
d) Poor curing
e) Cracking and microcracking
1 . Low-permeability concrete * achieved by:
a) Proper mixture proportioning †
b) Reduced unit water content
c) Increased cementitious material content
d) Appropriate use of supplementary cementitious materials (SCMs)
e) Air entrainment
f) Adequate consolidation
g) Effective curing ‡
2. Cracks and separations due to:
a) Loading/stress concentrations
b) Thermal stress
c) Shrinkage
2. Reduced tensile stress in concrete by: #
a) Using tensile reinforcement of adequate size, correctly located
b) Inclusion of pozzolan to reduce temperature rise
c) Provision of adequate contraction joints
d) Effective curing
3. Leaching and liquid penetration due to:
a) Flowing liquid §
b) Ponding
c) Hydraulic pressure
3. Structural design
a) Minimize areas of contact and turbulence
b) Provision of membranes and protective-barrier system(s) ||
c) P rovision o f adequate drainage and through- f ow
Factors that control permeability are discussed in more detail in Chapter 3.
The mixture proportions and initial mixing and processing of fresh concrete determine its homogeneity and density.
P oor c uring procedures
res ult in
f
aws and cracks .
Resistance to cracking depends on strength and strain capacity.
Movement o
f water- carrying
deleterious
subs tances inc reas es reactions that depend on both quantity and veloc ity o
f f ow.
Concrete that will be frequently exposed to chemicals known to produce rapid deterioration should be protected with a chemically-resistant protective-barrier system.
f
f the concrete, which can
further infltration o f aggressive
waters. However, brucite is not always observed in seawater
exposures. Magnesium can react with any hydrate normally
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present in concrete, including calcium silicate hydrate (C-S-H)
and hydrated magnesium silicate (M-S-H) phases, forming
phases such as hydrotalcite. Such phases can be detrimental
to concrete because they can reduce the binding capacity of
the cement paste. Sulfates in seawater may lead to the formation of secondary reaction products typically associated with
sulfate attack such as ettringite, gypsum, and thaumasite.
Chlorides may react with calcium aluminate phases to form
chloroaluminate phases such as Kuzel’s salt and Freidel’s salt.
Chlorides are also a particular concern because of the potential for corrosion of embedded steel.
Mather (1 966) noted that reactions of mature concrete with
sulfates in seawater is similar to that with sulfates in fresh
water or leached from soils, but the effects differ because
the presence of chloride ions alters the extent and nature
of the chemical reaction. Less expansion is found with a
cement of a given C 3 A content than would be expected of
the same cement in a nonmarine exposure where the water
has the same s ul fate ion content but lacks s ignifcant chlo ride concentration. Concretes made with portland cements
having C 3 A contents as high as 1 0 percent and subject to
continuous immersion in seawater have performed satisfactorily, provided that the permeability of the concrete is
low (Browne 1 980). USACE EM 11 1 0-2-2000:1 994 and
the Portland Cement Association recommend up to 1 0
percent calculated C 3 A for concrete that will be permanently
submerged in seawater if the w/cm is kept below 0.45 by
mass. Verbeck (1 968) and Regourd et al. (1 980) showed that
there may be a considerable difference between the calculated and measured clinker composition of cement, particumay fll pores on the outer sur ace o
reduce permeability and limit
larly with respect to the proportions of C 3 A and C 3 AF. Therefore, the interrelation between the measured C 3 A content and
the resistance to seawater may be uncertain. CSA A23.1 -1 4/
CSA A23.2 recommends the use of cements with between 4
Standard
Sharing Group and our chats
and 8 percent C 3 A unless slag or pozzolans are used.
7.2.3 Mitigation —The mitigation of seawater exposure requires the use of mixture proportions that minimize
permeability and tendency to microcracking, structural
designs that minimize the number and width of cracks, and
possibly the application of coatings that provide cathodic
protection or reduce permeability. The requirement for low
permeability is essential not only to delay the effects of
chemical attack on the cementitious phases in the concrete,
but also to afford adequate protection to reinforcement with
the minimum concrete cover recommended by ACI 357.1 R
for exposure to seawater. Steps to achieve low-permeability
concrete are discussed in more detail in Chapter 3 . Factors
such as w/cm , types of cementitious materials, appropriate
use of pozzolans and slag, aggregate grading, good consolidation, and adequate curing are important to achieve satisfactory performance.
The use of slag cement and silica fume may reduce the
permeability and increase the performance of concrete in
a marine environment (Lea 1 971 ; Fidjestøl and Frearson
1 994; Mather 1 981 a; Vanden Bosch 1 980). The permeability
of concrete made with appropriate amounts of suitable slag
cement or pozzolan can be lowered by orders of magnitude
compared to concrete of equal strength made with portland
cement only (Bakker 1 980). Concretes made with combinations of cement and silica fume, and with combinations of
cement, slag cement, and silica fume also have lower permeability and good performance in seawater exposure.
Concrete should be designed and constructed to minimize
the length, width, and number of cracks to limit seawater
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
access to the reinforcement. Additionally, concretes should
achieve an in-place strength of at least 4000 psi (27.5 MPa)
before being exposed to seawater. Marine structures often
involve thick sections and rather high cement factors. Such
concrete may need to be treated as mass concrete in which
the effect of the heat of hydration is considered. When this is
the case, the recommendations of ACI 207.1 R, ACI 207.2R,
and ACI 224R apply.
Conductive coatings applied at the time of construction
as part of a cathodic protection system may provide additional protection for concrete exposed to saline groundwater.
Coatings that s ignifcantly res trict the evaporation o f free
water from the interior of concrete can reduce resistance to
freezing and thawing.
7.3—Acid attack
contain s ulfde- bearing
minerals
s uch as pyrite that produce
sulfuric acid on oxidation. Further reactions can produce
sulfate salts, which may lead to sulfate attack (Hagerman
and Roosaar 1 955 ; Lossing 1 966; Bastiensen et al. 1 957;
Mourn and Rosenquist 1 959). Some mineral waters may
contain high concentrations of dissolved carbon dioxide and
hydrogen
s ulfde;
such
s olutions
can
be
acidic
and
highly
aggressive to concrete (RILEM 1 962; Thornton 1 978 ).
Organic acids may come from farm silage or from manufacturing and processing facilities such as breweries, dairies,
canneries, and wood-pulp mills. Animal feed and manure
may also contain acids that corrode concrete (De Belie et
al. 1 996). This can be of considerable concern in the case of
f
calcium phosphate and calcium oxalate, respectively. These
deposits are insoluble in water and tend to coat the concrete
surface, protecting against further deterioration. Exposure to
sulfuric acid may lead to rapid deterioration because of its
low pH. In addition, these reactions may produce calcium
sulfate, which may then drive sulfate attack of adjacent
concrete that was unaffected by the initial acid attack.
The decomposition of C-S-H by acid attack will typically produce a silica gel that has little binding capacity. The
decompos ition
products res ulting from the decalcifcation
of the original C-S-H have low solubility and can provide
some protection from further corrosion (Shi and Stegemann
2000; Shi 2003). If the original C-S-H has a high calcium/
silica ratio, more calcium will dissolve and the concrete will
corrode more quickly. Shi and Stegemann (2000) found that
a
In general, portland-cement concrete does not have good
resistance to acids, although some weak acids can be tolerated, particularly if the exposure is occasional.
7.3.1 Occurrence —The products of combustion of many
fossil fuels contain gases that can combine with moisture
to form acids . S ewage that is not s u ffciently aerated can
form sulfuric acid (Flemming 1 995 ; Sydney et al. 1 996).
Acids may occur in runoff from some mines and in some
industrial waters. Peat soils, clay soils, and alum shales may
oors , even where s tructural integrity is not impaired.
7.3.2 Reaction mechanisms —The deterioration of
concrete by acids is primarily the result of decomposition of
the hydration products of the cementitious paste. Different
cement hydration products start to decompose at different pH
values. Portlandite (CH, Ca(OH) 2 ) is the most soluble hydration product. At room temperature, portlandite decomposes
at pH below 1 2.4, ettringite decomposes at pH below 1 0.4
(Warren and Reardon 1 994), and C-S-H starts to decompose
when pH drops to around 1 0 (Beaudoin and Brown 1 992).
Aggregates made from limestone and dolomitic rocks are
susceptible to acid attack, while most siliceous aggregates
are resistant to acids.
The degree to which an acid is aggressive toward concrete
depends on the type of anion, its concentration, and its degree
of dissociation in the solution (Zivica and Bajza 2001 ). For
a given pH, acetic acid is more aggressive than nitric acid
(Shi and Stegemann 2000; Shi 2003 ; Bakharev et al. 2003 ).
Oxalic and phosphoric acids are less aggressive, primarily
because they react with concrete to form precipitates of
41
lime-
f
y
as h
pas te
corroded
more
slowly
than
portland
cement paste, although the former was more porous.
7.3.2.1 Carbonic acid attack—Carbon dioxide can
dissolve in rain to form carbonic acid, which may then enter
the ground. The decay of organic matter liberates carbon
dioxide that may also form carbonic acid in groundwater. The
concentration of carbonic acid in groundwater can become
high enough to attack concrete. Calcium carbonate dissolves
in the presence of carbonic acid to form free calcium (Ca 2+)
and bicarbonate (HCO 3 –) ions. Detailed discussions of
carbonate equilibria in natural systems can be found elsewhere (Stumm and Morgan 1 995 ; Krauskopf and Bird 1 995 ;
Butler 1 998 ). If the alkalinity of the soil is high enough, the
soil will neutralize or buffer the carbonic acid component of
the water, preventing carbonic acid attack of the concrete. If
the acid is not neutralized, it can attack concrete to varying
degrees, ranging from mild to s ignifcant. This type o f attack
has been referred to as aggressive CO 2 attack in the literature
when, in fact, it was not CO 2 but carbonic acid attack. Test
criteria relating to carbonic acid attack have used the term
“aggressive CO 2 ” (reported as milligrams of CO 2 per liter)
for what was really carbonic acid in s uffcient concentration
to attack concrete. Good discussions of this topic may be
found in Lea (1 971 ) and Hewlett (1 998).
Waters that have potentially harmful concentrations of
carbonic acid tend to have pH values ranging from approximately neutral to slightly acidic. Because the rate of attack
depends on both the properties of the concrete and concentration of the carbonic acid, neither the pH nor the amount
of free CO 2 in water is a reliable indicator of the degree
of potential harm. In addition, there is no consensus as to
the limiting values of pH or CO 2 concentration in water, in
part because of widely varying conditions in underground
construction. Studies have shown that water containing more
than 20 mg/L of carbonic acid (reported as mg/L of CO 2 ) can
result in rapid carbonation and attack of the hydrated cement
paste. However, freely moving waters with 1 0 mg/L or less
of carbonic acid (reported as mg/L of CO 2 ) can also result
in s ignifcant
carbonation
( Terzaghi
1 948 , 1 949; Hewlett
1 998 ). DIN 4030 includes both criteria and a test method
for assessing the potential of damage from carbonic-acidbearing water.
42
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
7.3.3 Mitigation —The use of pozzolanic materials such
f y ash and s ilica fume will decrease the CH content and
increase the resistance of concrete to acids (Sellevold and
Nilson 1 987). In all cases, however, exposure time to acids
should be minimized, if possible, and immersion should be
avoided. No hydraulic-cement concrete, regardless of its
composition, will long withstand highly acid water (pH of 3
or lower). Such cases typically call for the use of an appropriate protective-barrier system.
as
7.4—Fresh water
Fresh water refers to aqueous solutions with nearly neutral
pH, very low ionic strength, and low dissolved solids content.
7.4.1 Occurrence —Fresh waters include rainwater; waters
in most streams, rivers, and lakes; and domestic water that
is chlorinated and
f
uorinated.
F res h waters can als o occur in
industrial, manufacturing, and other facilities where distilled
waters are produced or used in various processes.
In nature, lightning produces weak nitrous, nitric, sulfurous, and sulfuric acids in natural waters that can cause some
surface deterioration of concrete, especially in areas that
experience frequent thunderstorms. Some fresh waters may
be somewhat acidic due to exposure to acid rain, or they may
contain small concentrations of sulfates, nitrates, and other
salts that, in higher concentrations, could attack concrete.
S ignifcant
chemical
attack
by
fresh water, however, is
virtually unreported. That concrete is not s ignifcantly dete riorated by fresh water is evidenced by highways, culverts,
Getaremore
from
pipes, and buildings that
built FREE
with thestandards
full expectation
that their function will not be s ignifcantly a ffected by s uch
exposure during their expected service life.
7.4.2 Reaction mechanisms —Very pure waters are aggressive because they are undersaturated with respect to the CH,
C-S-H, and calcium carbonate components of the cementitious paste. In general, the reaction mechanisms associated with freshwater exposures are similar to those of acid
attack because the attack involves preferential dissolution
or leaching of soluble cement hydration products. Consequently, constant replenishment of fresh water may accelerate
deterioration
and
f
owing
water may be
aggres s ive
as
well. Falling water from devices such as gutters and downspouts may be particularly aggressive because it introduces
a physical or erosive component to the attack.
Freshly cast concrete is highly alkaline and its surfaces
may be affected by exposure to fresh water. Most surfaces
carbonate readily when exposed to air, however, rendering
them largely stable. Unfortunately, many publications report
laboratory studies of fresh concrete samples with exposure
to chemicals be fore signifcant carbonation. S uch studies are
o ften inapplicable to feld concrete. S ome fresh water that is
undersaturated with respect to carbonates may attack carbonated concrete. Even under decades of such exposure, the deterioration of carbonated concrete is typically quite slow.
7.4.3 Mitigation —Strategies to minimize the effects of
fresh water exposure include minimizing permeability and
reducing the portlandite content of the cement paste. Design
considerations that provide adequate drainage, limit replenishment, and shelter against falling water are also important.
7.5—Carbonation
Carbonation occurs when hydrated cementitious
compounds react with atmospheric carbon dioxide or
carbonate ions in solution. The pore structure of concrete
largely determines the rate and depth to which carbonation
occurs (Bier 1 987). Carbonation begins at the exposed
surface of concrete to form an outer layer of carbonatebearing compounds, reducing the porosity of the surface
(Parrott 1 987). The reduction in porosity is directly related
to the conversion of CH to calcium carbonate, thus resulting
in an 11 percent increase in solid volume as compared with
the initial volume of CH within the upper layer (Bier 1 987).
Prolonged moist curing may delay carbonation (Parrott
1 98 7).
C arbonation
can
be
either
benefcial
or
harm ful,
depending on the age of concrete and the environment.
Although carbonation can improve the strength of concrete
and decrease permeability, it can also increase the rate of
corrosion of steel reinforcement.
7.5.1 Occurrence
7.5.1.1 Carbonation o f fresh concrete —The carbonation of fresh concrete occurs typically from exposure to
atmospheric carbon dioxide during the hardening process.
Carbonation may also occur due to use of unvented heaters
or the exhaust fumes of equipment. The result may be excessive surface cracking or a weak powdery residue (laitance)
on the s ur face. S evere carbonation prior to fnal s et can res ult
in a less wear-resistant surface.
7.5.1.2 Carbonation o f early-age concrete —Immature
Standard
and toourcarbonation
chats
concreteSharing
is more Group
susceptible
than mature
concrete
becaus e
its matrix has not hydrated s u ffciently
to limit permeability. Carbonation will, therefore, progress faster in the early history of a given concrete member.
Initially, CH will react with carbon dioxide to form calcium
carbonate; tetracalcium aluminate hydrate will react with
carbon dioxide to form monocarboaluminate hydrate (Bier
et al. 1 988).
7.5.1.3 Carbonation o fmature concrete —Aged concrete is
carbonated to some degree. Carbonation of concrete occurs
at expos ed s ur faces. More- permeable, coars e, open- fnis hed
concrete, and certain environments, however, can increase
the rate and depth of carbonation. Continued carbonation
can decrease the pH of the cementitious matrix, which leads
to an increased rate of corrosion of reinforcement. Refer to
ACI 222R for detailed discussion on corrosion.
7.5.2 Reaction mechanisms
7.5.2.1 Atmospheric carbonation —The reaction of hydrated
portland cement with atmospheric carbon dioxide is generally
slow. The rate is highly dependent on the relative humidity of
the environment, temperature, permeability of the concrete,
and the concentration of CO 2 . Keeping all other factors equal,
carbonation occurs most rapidly when the relative humidity of
the concrete is between 50 and 75 percent. Below 25 percent
relative humidity, the degree of carbonation that takes place is
insignifcant, and above 75 percent relative humidity, mois ture in the pores restricts CO 2 penetration (Verbeck 1 958).
7.5.2.2 Carbonation by carbon species in water—Carbonation can also take place when concrete is exposed to water
containing su ffcient concentrations o f carbonate or bicar-
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
bonate ions. Carbon species are found in most water. Carbon
species enter water when carbon dioxide in the air dissolves
in rain, as organic matter decays in soil, or as certain minerals
containing carbonates are weathered. Carbonate (CO 3 2–) and
bicarbonate (HCO 3 –) ions are usually the most abundant
carbon species found in natural waters. Detailed discussions
of carbonate equilibria in natural systems can be found in
Stumm and Morgan (1 995), Krauskopf and Bird (1 995), and
Butler (1 998).
7.5.2.2.1 Alkalinity o fwater (carbonate alkalinity bu ffer) —
Carbonate alkalinity can play a critical role in the service life
of concrete. In general, when carbonate alkalinity, which is a
concentration o f carbonate and bicarbonate ions , is s u ffcient
to neutralize or buffer the acidic component of water, both
acid attack and leaching may be prevented. Leaching attack
can occur when water is consistently low in carbonate alkalinity, causing the selective dissolution of carbonate-bearing
compounds of cement. This leaching attack, referred to as
aggressive CO 2 attack in past literature (Biczok 1 967; Lea
1 971 ; Ibrahim et al. 1 997), is better termed “low-carbonate
alkalinity attack”.
Where free CO 2 attack, or carbonic acid in solution,
relates to acidic attack that dissolves concrete from the
surface inward (4.3 ), low-carbonate alkalinity attack relates
to leaching attack from within concrete. Simply stated, lowcarbonate alkalinity is an imbalance in water concerning the
lack of carbonate alkalinity needed to neutralize or buffer
the acidic component, and low-carbonate alkalinity attack,
or the term aggressive CO 2 attack, is the leaching of the
cementitious carbonate-bearing compounds until an equilibrium has been reestablished. This guide does not detail this
complicated chemistry, but rather explains the importance of
carbonate alkalinity in most placement environments where
water is in direct contact with concrete. Discussions of the
carbonate alkalinity buffer system in concrete are found in
Lea (1 971 ), Biczok (1 967), and Hewlett (1 998).
7.6—Industrial chemicals
Industrial chemicals may attack concrete in different
depending on the type or clas s ifcation o f the chem ical, its concentration, the duration of exposure, and interactions with the components in the concrete. Industrial
processes usually result in exposure conditions ranging from
incidental to continuous, and even extreme, when industrial
process factors such as high temperatures, high humidity,
and equipment vibration exacerbate the situation.
7.6.1 Occurrence —Industrial chemicals may include
acids, bases, alkalis, corrosives, oxidizers, combustibles,
ways ,
f
ammables ,
explos ives ,
cryogenic,
and
other
process -
f thes e conditions .
Refer to PCA (2001 ) and ACI 51 5.2R for summaries of the
effects of many industrial chemicals on concrete.
Concrete in industrial settings may experience direct
exposure to chemicals, indirect exposure to chemicals, and
exposures where abrasion and erosion are a concern. Direct
exposures occur in structures such as those handling cooling
water and primary containment structures. Abrasion and
erosion must be considered in structures subject to moving
s pecifc
conditions
or
combinations
o
43
liquids. Foundations, equipment supports, and structural
framing may experience direct contact with chemicals but
are more frequently subject to indirect chemical exposure.
Indirect exposure includes fumes or vapors and precipitants
from these sources.
Anhydrous chemicals may or may not be aggressive
to concrete in their pure form but can become aggressive
when exposed to moisture, such as humidity in the air. Upon
contact with humidity, these compounds may deliquesce and
then infltrate the s urrounding concrete.
7.6.2 Reaction mechanisms —Because industrial chemicals range widely from highly alkaline to very acidic, a
detailed discussion of the potential reaction mechanisms is
beyond the scope of this guide. Biczok (1 967) gives useful
background information on a wide range of these chemicals.
Alkaline solutions have a pH greater than 7. Because
concrete is highly alkaline itself, the interaction of alkaline
chemicals from industrial processes may not affect the durability of the concrete directly unless there is a component
or characteristic in the concrete structure that would react
with the penetrating chemical. An example is the corrosive
action of salt solutions on reinforcing steel in the concrete.
Leaching by more neutral solutions can increase the porosity
of concrete, making it more susceptible to freezing-andthawing damage, spalling, and further attack. Precipitant
deposits could plug air entrainment spaces, making the
concrete more susceptible to freezing-and-thawing damage
and spalling. Acids and corrosives do not penetrate the
concrete to any appreciable depth but react on contact with
the concrete surface via the mechanisms described in 7.3.
7.6.3 Mitigation —All the factors that apply to creating
durable concrete apply to durability against chemical attack.
Permeability of the concrete can be decreased by supplementary cementitious materials (SCMs) in the design mixture
in conjunction with good curing. Reinforcing steel protection can be achieved by several methods, including sealing
the concrete, removing any surface contamination from the
reinforcement, improving the bond between the reinforcement and the paste, using epoxy-coated reinforcement,
and increasing the minimum cover over the reinforcement.
When increasing the reinforcement, the architect/engineer
must consider that crack widths will likely increase. Chapter
3 discusses these topics in detail.
There are other options available to the design engineer
and the contractor that will increase concrete durability. Joint
fllers, j oint sealers, waterstops, and surface sealers that are
resistant to and compatible with the chemicals that the structure is expected to be exposed to should be specifed. Every thing should be installed or applied in full compliance with
manufacturers’ recommendations. Wet curing is the recommended method to achieve the most durable concrete. Curing
agents, if used, must be compatible with the sealer or surface
coating that will be applied. Using a shrinkage-compensating
concrete in accordance with ACI 223R can improve durability,
particularly in a chemical environment. Because shrinkagecompensating concrete minimizes shrinkage cracking and the
number of contraction joints, it reduces the potential leakage
paths in an industrial environment.
44
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
7.7—Deicing and anti-icing chemicals
The use of chemical agents has a long history in helping
maintain safe winter driving conditions. Such maintenance
generally involves two different strategies: deicing and antiicing. Deicing refers to the removal of ice after deposition on
a pavement, whereas anti-icing refers to the prior application
of chemicals to prevent the adherence of ice to the pavement. Historically, most durability problems associated with
deicers were linked to physical processes that exacerbate
scaling, rather than chemical attack. However, deicers may
be associated with signifcant deterioration from various
chemical attack mechanisms.
7. 7. 1 Occurrence —Deicers fall into two broad groups:
chloride-based and non-chloride-based. Chloride-based
solutions have seen more widespread use historically.
They include sodium chloride (NaCl), calcium chloride
(CaCl 2 ), magnesium chloride (MgCl 2 ), and many commercially available products comprising combinations of these
salts. Some chloride-based deicers may contain signifcant
concentrations of other chemicals. Other product formulations include MgCl 2 -based agricultural products. The use
of MgCl 2 deicers has increased signifcantly since 2000,
particularly in the western United States. In part, this growth
resulted because MgCl 2 is more effective at lower temperatures than NaCl and CaCl 2 .
The introduction of nonchloride deicers stemmed from
concerns of corrosion of reinforcing steel and environmental
impacts on vegetation, and because they are also more effecGet than
more
FREE standards
tive at lower temperatures
chloride-based
deicers. from
The
principal non-chloride deicers include calcium-magnesium
acetate; urea; glycols consisting of ethylene, propylene, and
diethylene glycols; and alkali-acetates and alkali-formates.
Calcium-magnesium acetate deicers were introduced in the
late 1 970s. Their use has become more widespread with
decreasing production costs. Urea was traditionally used
for airfeld pavements but is less commonly used now due
to environmental concerns. Ethylene and propylene-based
glycols are in widespread use, primarily for deicing and antiicing of aircraft. They are also occasionally used for maintenance o f airfeld pavements, sometimes in combination with
urea. A new generation of alkali-acetate and alkali-formate
deicers and anti-icers emerged in the late 1 980s and early
1 990s to replace urea in deicing airfeld pavements and to
alleviate concerns associated with the toxicity of ethylene
glycol. These deicers and anti-icers include potassium
acetate, sodium acetate, potassium formate, and sodium
formate, and are widely used.
7. 7. 2
Reaction mechanisms —Reaction mechanisms
linked to deicing chemicals are complex and form an active
area of ongoing research. Although initially regarded as
benign for concrete, NaCl is now understood to drive reactions that can lead to portlandite dissolution. The dissolution of portlandite may increase the porosity of the concrete
and lower the pH of the pore solution, which may destabilize the C-S-H phase. Chloride-based deicers may drive
the formation of complex calcium chloroaluminate phases
such as Freidel’s salt. Numerous studies show that CaCl 2 is
aggressive to concrete (Collepardi et al. 1 994). Among other
mechanisms, CaCl 2 deicers drive reactions that may form
hydrated calcium oxychloride phases (Brown and Bothe
2004). The generation of hydraulic pressures from these
reactions may be disruptive to the cementitious paste. Chloride may also accelerate alkali-silica reaction (ASR) under
some conditions (Chatterji et al. 1 986). MgCl 2 deicers are
linked to the formation of brucite, which is not damaging,
and magnesium silicate hydrate phases (M-S-H) that form
at the expense of C-S-H. The formation of M-S-H, which
is not cementitious, can produce signifcant deterioration in
pavements by cracking, delamination, and, ultimately, disintegration. Disruptive oxychlorides have also been found in
mortars exposed to MgCl 2 (Julio-Betancourt and Hooton
2005 ; Sutter et al. 2006).
Calcium-magnesium acetate may be among the most
aggressive deicers in terms of chemical attack. Some authors
(Peterson 1 995 ; Santagata and Collepardi 2000) report that
exposure to calcium-magnesium acetate deicer solutions
signifcantly degrades the cement matrix, resulting in loss
of mass and compressive strength. The reaction mechanisms are similar to those of MgCl 2 deicers—dissolution
and leaching of portlandite, destabilization of C-S-H and
formation of M-S-H, and precipitation of brucite and calcite.
Calcium-magnesium acetate deicers are linked to scaling,
cracking, and loss of mass and compressive strength.
Premature deterioration o f some airfeld concrete pavements
exposed to alkali acetate and sodium formate deicers has
caused concern. Research is ongoing to investigate potential
Standard
Sharing
Group
and
links between
these
deicers
andour
the chats
durability of concrete.
Rangaraju et al. (2005) suggested that alkali-acetate and
alkali-formate deicers may cause deleterious expansions due
to ASR in test specimens.
Chemical deicers may also contribute to the relativ ely rapid
deterioration o f joints in pavements and exterior f atwork.
The damage manifests as cracking and spalling parallel to
joints that may be most severe at joint intersections and in the
wheel path. Some states in the northern United States have
observed premature deterioration at pavement joints (Taylor
et al. 201 2). The relationship between chemical deicers and
accelerated joint deterioration is an area of active research
and several mechanisms are proposed to explain this deterioration. In addition to depressing the freezing point of water,
deicing and anti-icing chemicals induce fundamental changes
in the physical and chemical properties o f solutions that fll
joints, such as their viscosity, surface tension, and sorption
(Spragg et al. 2011 ; Villani et al. 201 4a). These changes
result in a higher degree of saturation and marked increase
in the frequency of cracking and microcracking events under
certain temperature cycling conditions (Villani et al. 201 4b;
Farnam et al. 201 4). Concrete exposed to high concentrations o f chemical deicers tends to have air voids flled with
secondary deposits that include ettringite, portlandite, and
oxychloride minerals (Sutter et al. 2006; Peterson et al.
201 3 ). Oxychloride phases are commonly observed in laboratory-based studies, but documenting their presence in feld
concrete remains di ffcult due to their instability (Peterson et
al. 201 3). Ettringite deposits may accelerate the saturation of
concrete (Stark and Bollmann 1 999) and possibly diminish
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
the e ffcacy o f air void s ystems , making the concrete more
susceptible to damage during freezing-and-thawing cycles.
Deicing salts may also preferentially dissolve calcium-based
compounds at low temperatures in the interfacial transition
zone of coarse aggregate particles exposed in the saw cut
(Zhang and Taylor 201 2). Some cracking has been detected
in concrete exposed in the laboratory to deicer solutions
at temperatures above freezing (Farnam et al. 201 4). This
cracking may occur as a result of mineralogical changes in
the cement paste that result from reactions involving deicing
solutions , although the s pecifc reaction mechanis ms remain
uncertain at this time.
7.7.3 Mitigation —Mitigation of the effects of chemical
deicers includes steps described previously to minimize
the permeability and control the reactivity of the concrete.
Good curing is especially important in mitigating the effects
of deicers. Care must be taken to avoid exposure to deicers
during the frs t year o f s ervice. Mitigation als o requires the
provision and maintenance of adequate drainage to minimize duration of exposure.
7.8—Environmental structures
Environmental structures are designed to contain liquids
and gases. Environmental structures include water treatment
plants; domestic and industrial wastewater treatment plants;
storage tanks and reservoirs; water and wastewater pump
stations; conduits, sewers, manholes, and junction chambers;
and
hazardous
materials
containment
s tructures
defned
in
ACI 350.2R. When the concentration of the contained chemicals is su ffcient, they can attack the concrete. The chemi cals that are contained may be bulk process chemicals or
those that exist in the liquid or gas contained in the environmental structures. The attack can be more aggressive at
higher temperatures.
7.8.1 Occurrence —Aggressive chemicals typically
occur in bulk storage containers and in pumping systems.
Secondary containment structures must keep spilled chemicals from reaching the soil. Chemical attack of the concrete
may occur at the point of injection of the chemicals into the
process system, particularly if the injection is toward the
concrete rather than into the process liquid.
Chemical agents in environmental structures range widely
from relatively benign to highly aggressive in terms of their
ability to attack concrete. ACI 350 class ifes chemical agents
in three broad categories. Refer to R4.5.1 .4 of ACI 350.1
for a more complete discussion. Group 1 chemicals are not
considered to be directly harmful to concrete but may be a
concern if they combine with other chemicals that can react
with concrete; Group 2 chemicals such as activated carbon
and potassium permanganate may stain concrete; and Group
3 chemicals are corrosive to concrete. ACI 350.1 differentiates Group 3 chemicals according to the rate at which they
will corrode concrete under typical exposure conditions
associated with environmental structures. Group 3a chemicals have a slow rate of corrosion, Group 3b chemicals have
a moderate rate of corrosion, and Group 3c chemicals have a
rapid rate of corrosion.
45
7.8.2 Reaction mechanisms —The reaction mechanisms
that attend environmental structures range widely, as there
are a multitude of different types of exposure conditions.
In general, the reaction mechanisms of many exposures
are similar to many of the mechanisms discussed in other
sections of this chapter. Environmental structures exposed
to chemicals with low pH are subject to leaching and corrosion of the concrete due to the dissolution of cementitious
components, and possibly aggregates, in the concrete. Some
aggressive chemicals may react with cementitious phases
to stain the concrete. Water treatment plants may expose
concretes to relatively fresh water. Some structures may
encounter exposure to solutions that have very low alkalinity.
Wastewater treatment plants and facilities may expose
concrete to both chemical and biological activity, resulting in
highly aggressive exposure conditions. Bacterial action under
anaerobic conditions may lead to the generation of hydrogen
sulfde gas. Aerobic bacteria present on wet sewage facility
walls, such as pipe linings, may convert hydrogen sulfde into
furic acid. In these environments, signifcant deterioration
of the concrete can occur from the dissolution of cementitious phases such as portlandite and C-S-H, the deposition of
secondary products such as gypsum, and leaching of ferruginous and calcareous components in the aggregates. Under
other conditions, solutions in wastewater treatment plants
lead to the deposition of struvite, which is a magnesium
ammonium phosphate mineral (MgNH 4PO 4· 6H 2 O). Struvite
deposits can impede the e ffciency o f treatment processes and
cause maintenance problems.
7.8.3 Mitigation —The process design of environmental
concrete structures can affect the potential for chemical
attack of the concrete. ACI 350 establishes the minimum
requirements for the design of concrete environmental structures, including s pecifc durability and protection require ments. Chemical attack in environmental structures may be
controlled by using a process and structural design that does
not increase the corrosiveness of the liquid being processed;
use of a properly designed concrete mixture for the expected
service conditions; use of properly selected concrete materials; and use of good concreting procedures related to
handling, placing, and curing. Even when these procedures
are followed, it may still be necessary to provide a protective
barrier, especially when the chemical that may be in contact
with the concrete can cause unacceptably rapid deterioration
for the expected time the concrete must contain the chemical. Refer to ACI 350 for requirements concerning the use
of protective barriers.
sul
CHAPTER 8—PHYSICAL SALT ATTACK
8.1 —Introduction
Physical salt attack is a deterioration mechanism caused
by crystallization of salts in pores near concrete evaporative surfaces. The mechanism is called physical salt attack
because chemical reactions between concrete and crystallizing salt are not involved (Mehta and Monteiro 2006).
D eterioration ranges from very fne s ur face crumbling and
scaling, which is primarily cosmetic, to severe progressive
46
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
disintegration. Deterioration occurs at evaporative surfaces
ab o ve
the
soil
li ne
or
on
f
atwo rk
whe re
water
may
wic k
(Fig. 8.1 a and 8.1 b). The most common salts linked to
physical salt attack include, in order of decreasing aggressiveness, sodium sulfate, sodium carbonate, and sodium
chloride.
Physical salt attack has also been called salt crystallization, salt hydration distress, salt damp, and salt weathering.
Reports identifying physical salt attack include Bates et al.
(1 91 3), Wig et al. (1 91 7), Reading (1 975 , 1 982), Novak
and Colville (1 989), Yen and Bright (1 990), Haynes et
al. (1 996), Hime et al. (2001 ), and Erlin and Jana (2003).
Physical salt attack on concrete began to receive attention
in the early 1 990s, with research initiated by Folliard and
Sandberg (1 994) and theoretical developments by Scherer
(1 999). Prior research generally focused on chemical sulfate
attack on concrete, while physical sulfate attack was often
overlooked or mis identifed.
Physical salt attack also occurs as a deterioration mechanism on exposed stone and brick masonry, as well as other
porous materials. This deterioration mechanism is an important component of the broader phenomenon known as salt
weathering, where differential expansion and contraction of
pore solutions within porous building materials plays a role
in deterioration. Literature from the geosciences, as well as
building and art conservation, provides useful information
on the phenomena associated with salt weathering (Doehne
2002; Evans 1 970; Goudie and Viles 1 997; Winkler 1 997).
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standards
from
Scaling of concreteGet
surfaces
by physical
salt attack
should not be confused with scaling of concrete surfaces by
freezing and thawing of concrete in the presence or absence
of deicing salts. Concrete surfaces exposed to freezing and
thawing conditions in the presence of deicing salts may deteriorate by chemical and physical mechanisms (Chapter 4)
(Mehta and Monteiro 2006; Valenza and Scherer 2007).
Fig. 8.1a—Physical salt attack on concrete residential foun dation. Scaling o f the concrete surface resulted from sodium
sulfate in the soil pore water. The salt accumulated behind a
paint coating on the stem wall (Haynes and Bassuoni 2011).
Standard Sharing Group and our chats
8.2—Occurrence
Physical salt attack occurs throughout the world. The
process requires water-soluble salts from seawater, groundwater, soil, and other sources and ambient environmental
conditions , which us ually involve f uctuations in tempera ture and relative humidity, typically on a diurnal basis. In
natural environments, this commonly occurs in arid regions,
such as the southwestern United States; portions of southern
Europe; coastal areas of Australia; and in much of the Middle
East, particularly in the Gulf regions. Additionally, localized
microclimates can provide similar conditions. Landscaping
practices and irrigation can also lead to conditions conducive to physical salt attack.
Physical salt attack may occur even when salt concentrations in soils are low because the salts concentrate over time
at concrete evaporative surfaces. Unlike chemical sulfate
attack, there is no reported threshold concentration of salts
or chemical components that indicate the potential severity
of attack. A useful approach for assessing the potential
for attack involves understanding the potential for cyclic
wetting and drying of evaporative surfaces, and determining
the types of water-soluble salts in contacting soils and water
Fig. 8.1b—Physical salt attack on concrete garage slab
caused by sodium carbonate. Scaling occurred along an
evaporation front (Haynes and Bassuoni 2011).
(atmospheric, ground, and irrigation). This involves determining whether soils or water contain water-soluble anions
such as sulfate, carbonate, bicarbonate, and chloride, and
cations such as sodium, calcium, potassium, and magnesium.
Solutions of sodium sulfate contain sulfate ions, so the
potential for chemical sulfate attack on concrete exists
when this salt is present in the exposure environment. Other
sulfate salts, such as calcium, magnesium, potassium, iron,
and ammonia sulfate, can participate in chemical sulfate
attack on concrete, but do not appear to damage concrete by
physical salt attack.
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
8.3—Background
Much of the early literature reported evidence of physical
salt attack by observations o f feld-exposed concrete, usually
in sulfate-bearing soils and occasionally in carbonatebearing soils. In 1 91 2, tests were conducted by Bates et al.
(1 91 3) to investigate the cause o f feld distress to concrete
exposed to alkali waters (mostly sodium sulfate). This led to
a major test program incorporating test sites in seven areas of
the United States known to have soils with high concentrations of alkali salts. Williams and Furlong (1 926) observed
that crystallization of salts in concrete pores occurred along
with some other unknown chemical action. The U.S. Bureau
of Reclamation (USBR) (1 963) noted that salts such as
sodium carbonate can cause surface disintegration by crystallizing in concrete pores, and that this action appears to be
purely physical. Later, Reading (1 975, 1 982) reported that a
concrete tailrace wall of a dam showed deterioration caused
by physical action of sodium sulfate.
One o f the more signifcant test programs identi fying phys ical salt attack was the Portland Cement Association’s longterm feld tests on concrete exposed to sul fate-containing soils
(McMillan et al. 1 949; Stark 1 982, 1 989b, 2002). Overall,
thousands of specimens were tested. The early work by
McMillan et al. (1 949) identifed and discussed the anhy drous form (thenardite [Na2 SO 4]) and the decahydrate form
(mirabilite [Na2SO 4· 1 0H 2O]) of sodium sulfate, but these salt
phases were not associated with a deterioration mechanism.
Stark (2002) examined deteriorated specimens microscopically and identifed physical salt attack as a main cause o f
deterioration in the feld-exposed concrete specimens. Test
results showed resistance to deterioration was improved with
decreasing w/cm for concrete mixtures with and without
supplementary cementitious materials (SCMs). Concrete
mixtures found most resistant to physical salt attack contained
portland cement with no SCMs and 0.38 w/cm .
Stark (1 982, 1 989b, 2002) found that SCMs added to
concrete may decrease the resistance to surface deterioration.
Other investigators also observed detrimental effects related
to SCMs (Irassar et al. 1 996; Bassuoni and Nehdi 2009).
Folliard and Sandberg (1 994) conducted exploratory tests
on small concrete specimens to identify mechanisms of deterioration by inducing salt crystallization due to hydration,
evaporation, and changes in temperature. The most aggressive test environment was that of submerging specimens in
sodium sulfate solution and cycling the temperature between
41 and 86°F (5 and 30°C), with damage occurring as the
temperature decreased. The salt solution became supersaturated with respect to mirabilite as the temperature decreased;
hence, mirabilite precipitation or growth was likely responsible for deterioration. Haynes et al. (2008 , 201 0) reported on
a 3-year study in which concrete was partially submerged in
sodium sulfate, sodium carbonate, or sodium chloride solution and found that cycling of ambient environmental conditions caused more deterioration than when ambient conditions remained steady; however, damage was still observed
even when the ambient conditions were held steady. Sodium
sulfate was the most aggressive salt, followed closely by
sodium carbonate. Sodium chloride exposure produced
47
minor deterioration in laboratory testing in comparison to
the other salts.
8.4—Mechanism
Crystallization pressure is the primary cause of physical salt
attack. Scherer (2004a) provides a summary of the equilibrium thermodynamics that govern the development of crystallization pressure. As a general rule, crystallization pressures
increase with decreasing pore size within the concrete. These
pressures impose stresses on pore walls that ultimately cause
microcracking when pressures exceed tensile strength.
Salt crystallization occurs at evaporative surfaces because
the salts concentrate with evaporation, the solution becomes
supersaturated, and the salts precipitate. Once the salt crystallizes, they grow and generate pressure (Scherer 2004b).
Cycles of dissolving and recrystallization also cause damage.
For example, sodium chloride is hygroscopic; it absorbs
water from the air at relative humidities above 75.5 percent
(temperature range of 32 to 86°F [0 to 30°C]) until the crystals
dissolve, and then recrystallizes at lower relative humidity.
Sodium sulfate and sodium carbonate dissolve and recrystallize by a different mechanism. They can experience phase
changes due to changes in ambient temperature and relative humidity. Thenardite (Na 2 SO 4 ), the anhydrous form
of sodium sulfate, is stable at ambient conditions of 68°F
(20°C) and relative humidity up to 75 percent. At higher
relative humidity, however, the crystals will absorb moisture from the air and dissolve, from which the decahydrate
form (mirabilite [Na2 SO 4 · 1 0H 2 O]) crystallizes. Above a
temperature of 90.3°F (32.4°C), only thenardite is stable.
These characteristics of the salt permit phase change to occur
on diurnal cycles, with thenardite present during the hot, dry
daytimes and mirabilite during the cold, damp nighttimes. The
phase changes are accompanied by a change in crystal size.
Thenardite-to-mirabilite conversion is accompanied by a 31 4
percent increase in crystal size. For sodium carbonate, conversion of the low-hydrate phase (thermonatrite [Na2 CO 3 · H2 O])
to the decahydrate phase (natron [Na2CO 3 · 1 0H 2 O]) is accompanied by a 260 percent increase.
Factors important to distress by crystallization pressures
are the degree of supersaturation of the solution, the distribution of the salt in the pores (proximity of the evaporation
front to the surface), pore size and pore size distribution,
and sorptivity and tensile strength of concrete. The degree
of supersaturation is important because the potential for
crystallization increases with supersaturation and can lead
to more damage. Thermodynamic models show that higher
crystallization pressures develop in smaller pores (Scherer
2004a). Consequently, e ff orescence is usually not linked to
salt-based deterioration because salt crystallization occurs
on exposed surfaces where such growth is easily accommodated. Evaporation fronts, however, can develop below
exposed surfaces when fne pore networks decrease the rate
of transport of solution to surfaces, which can also happen
when evaporation is rapid. Where supersaturation occurs
below exposed surfaces, crystallization occurs within the
small pores rather than on free surfaces. This is known as
sub f oresence rather than e ff orescence. When sub f oresence
48
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
occurs, damage may result because pressures develop in the
fne pores, which can cause microcracking.
8.5—Recommendations
Research directed to better understand this mechanism of
deterioration is needed, particularly in the areas o f defning the
salt concentrations that can lead to deterioration, determining
the ambient environmental conditions that cause various
degrees of deterioration, and understanding the properties of
concrete that make it susceptible to physical salt attack.
Specifc recommendations cannot be made to prevent phys ical salt attack; however, the salts of sodium sulfate and sodium
carbonate are primarily responsible for physical salt attack on
concrete, while sodium chloride causes less deterioration. Other
common salts such as calcium sulfate and magnesium sulfate
do not participate in physical salt attack. While Stark’s longterm study on concrete exposed to sulfate soils is the main reference work on the topic, the test program inadvertently obtained
results on physical salt attack. Two trends were observed:
concrete having low w/cm showed less deterioration than
concrete having high w/cm , and concrete mixtures containing
fy ash or slag cement did not perform as well as companion
specimens containing only portland cement.
Where the risk of physical salt attack is unacceptable,
construction methods should be used that separate concrete
from contact with salt solutions, such as using capillary
breaks or protective coatings.
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CHAPTER 9—CORROSION
OF standards
METALS from
AND DEGRADATION OF OTHER MATERIALS
EMBEDDED IN CONCRETE
9.1 —Introduction
Understanding the conditions that cause corrosion
(rusting) of reinforcing and prestressing steel is vital. The
risk of excessive corrosion in concrete structures containing
embedded steel can be minimized to promote long service
lives. The purpose of this chapter is to summarize the mechanisms of steel corrosion, the conditions under which such
corrosion occurs, the methods and techniques that can be
used to prevent or limit steel corrosion, and the preservation
of other embedded materials.
Concrete protects against corrosion of embedded steel
because of the highly alkaline environment provided by the
pore f uid o f the portland cement paste. The adequacy o f
the protection depends on the depth of concrete cover, the
quality of the concrete, the details of the construction, the
degree of exposure to chlorides from concrete component
materials and from the environment, and the service environment. A more comprehensive treatment of the subject can
be found in ACI 222R and Broomfeld (2007) .
9.2—General principles o f corrosion initiation in
concrete
9.2.1 General —The process of corrosion of steel in
concrete is divided into several phases:
1 ) Initiation: the normal protective passive layer on the
steel breaks down
2) Corrosion growth (propagation): the (active) corrosion
process is established and corrosion progresses
3) Damage: corrosion is suffciently severe that cracking,
spalling, or both, occur and eventually the structural element
may not perform its intended function.
9.2.2 Protection mechanism in concrete
9.2.2.1 General —As described in ACI 222R, the high
alkalinity, with a pH greater than 1 2.5, of concrete protects
embedded steel reinforcement in concrete from corrosion.
When oxygen is present, the high pH of the pore solution
causes an ultra-thin corrosion flm to form on the steel
surface, termed a “passive flm”. The composition o f this
flm depends upon the metallurgy o f the metal and is understood to be a combination of hydroxides and oxides. This
flm is in equilibrium with the environment, slows corro sion reactions, and, thus, the steel is protected against active
corrosion and is said to be “passivated”.
Depending on the penetrability of concrete cover over
the steel and the alkalinity of the concrete pore solution, the
passive flm is maintained. I f the passive flm breaks down,
termed “depassivation,” corrosion rate accelerates and the
propagation phase begins. The flm can break down locally
so that localized corrosion results. If breakdown occurs over
larger areas, more uniform general corrosion takes place.
The primary causes o f flm breakdown include:
a) Chemical, physical, or mechanical degradation of the
concrete cover
b) Chloride penetration to the reinforcement
Standard
Sharing of
Group
and our
chats
c) Carbonation
the concrete
to reinforcement
depth
d) Change of polarization of the reinforcing steel such as
in dissimilar metal corrosion or stray current leakage.
9.2.2.2 Corrosion process —Corrosion of steel in concrete
is an electrochemical process that requires the development
of an anode where oxidation takes place and a cathode where
reduction takes place. At the anode, electrons are liberated
and ferrous ions are formed.
Fe → Fe ++ + 2e –
(9.2.2.2a)
At the cathode, electrons are consumed and hydroxyl ions
are formed and liberated.
2H 2 O + O 2 + 4e – → 4(OH) –
(9.2.2.2b)
The ferrous ions may subsequently combine with oxygen
or hydroxyl ions and produce various forms of corrosion
products or rust. The formation of rust often causes expansion that, in turn, may cause cracking and spalling of the
concrete cover. Refer to ACI 222R for a more detailed
description of the corrosion process.
9.2.2.3 Breakdown due to insu ffcient oxygen supply —
The passive layer requires an oxygen f ux corresponding to
approximately 0.2 to 0.3 mA/m 2 (1 .3 × 1 0 –4 to 1 .9 × 1 0 –4
mA/in. 2 ). I f the oxygen f ux is less than this, the passive flm
will gradually reduce in thickness, exposing bare steel. The
result is corrosion at an extremely low rate (corresponding
to the oxygen f ux), but at an active, though very negative,
potential (Fidjestøl and Nilsen 1 980). Such low oxygen
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
diffusion can occur in submerged concrete, in seawater or
fresh water (Fidjestøl et al. 1 985 ), or by exposure below the
groundwater table.
9.2.2.4 Carbonation —The reaction between CO 2 ,
commonly from the ambient air, and cement paste can reduce
the pH to less than 9, resulting in the loss of passivity. Corrosion rate increases, and the rate depends on availability of
oxygen as well as the moisture content, because this affects
the electrical resistivity of the concrete.
9.2.2.5 Local breakdown due to chloride —Chloride above
a certain concentration known as the chloride threshold will
cause local breakdown of the passive layer, leading to corrosion. The rate of corrosion depends on the availability of
oxygen and chlorides, the anode-cathode area ratio, and the
electrical resistivity of the concrete. The chloride concentration necess ary to cause breakdown o f the flm depends on
the pH and composition of the pore water, the quality of the
oxide flm, and the characteris tics o f the s teel and concrete
interface, and is therefore not one value for all cases. Note
that the chloride threshold is a distribution of values.
Although modern concrete has more variations in poref uid composition, chie f y due to the us e o f s upplementary
cementitious
materials
(S CMs )
such
as
f
y
as h,
s lag,
and
silica fume, the limits for chlorides in new construction as
given in ACI 222R are still conservative and appropriate.
ACI 31 8-1 4 contains higher allowable limits.
Because some concrete materials contain chloride that
will not be released into the concrete, past good performance
of these materials may provide a basis for permitting higher
chloride contents. The suggested chloride contents provide
a conservative approach that should result in low risk of
corrosion; this conservative approach is necessary because
o f con f icting reports on chloride thres holds , the e ffects o f
different exposure environments, and materials combinations. The conservative approach is also recommended
because exposure conditions, such as those encountered in
bridge decks, parking structures, and marine environments,
allow the penetration of chlorides from the environment.
Concrete should be made with constituents such that the
total chloride content in the concrete is within the guidelines
given in ACI 222R.
9.3—Propagation o f corrosion
9.3.1 General —Once corrosion has been initiated, a structure may still have many years of service life, especially if
the rate of corrosion is very low. The factors in 9.3.2 through
9.3.4 control the corrosion rate.
9.3.2 Anodic control —Anodic control is bas ed on s uff ciently controlling the rate of dissolution of corrosion products formed at the anode. The rate at which the dissolution of
iron takes place determines the corrosion rate.
9.3.3 Cathodic control—The rate of corrosion is
controlled by the availability of oxygen at the cathode and
the ratio between cathodic and anodic areas, and is limited
by the availability of oxygen, the size of the cathode, or both.
Normally, the supply of oxygen at the cathode far exceeds
that needed to sustain corrosion, so the rate is controlled
by other factors. Coating the reinforcement is one way
49
of limiting the oxygen supply to the cathode surface. The
coating also prevents access of aggressive media to the steel
surface. Cathodic control also occurs where the concrete
is completely water-saturated, which greatly reduces the
oxygen f ux from the concrete s ur face to the s teel.
9.3.4 Resistivity control — An electrolyte is essential
for corrosion to propagate. Resistivity of the electrolyte
can limit corrosion propagation for uncracked concrete
(Marcotte and Hansson 2003 ). Resistivity control requires
the electrical resistance of the concrete to be high enough
as to limit the current that can be developed from the two
half-cell reactions.
There may be a large difference in half-cell potential
between the cathodic and anodic areas of the steel. If the
electrical resistance R of the concrete between the two areas
is s u ffciently large, however, most o f the potential di fference is spent in overcoming the voltage drop IR , caused
by the current f ow I against this resistance, even at minute
corrosion current densities. This effect can often be seen in
fully carbonated concrete in a reasonably dry environment,
in concrete that contains SCMs, or where there is a great
distance between anodic and cathodic areas. This means that
corros ion can be ins ignifcant des pite large di fferences in
half-cell potential.
9.4—Corrosion-related properties o f concreting
materials
Materials for concrete should satisfy relevant standards
for structural concrete.
9.4.1 Portland cement —The alkalinity of portland cement
paste results from the presence of hydroxides of calcium,
potassium, and sodium in the pore solution. Calcium
hydroxide (CH) is the most abundant, and constitutes 1 5
to 25 percent of the paste. While the pH of saturated solutions of CH is only 1 2.4, the pH of 1 3.5 to 1 4 often found in
concrete pore water (Justnes and Nygaard 1 994) is explained
by the OH – ions associated with alkalis in the concrete.
The presence of C 3 A in the cement can have two benefts :
reducing both chloride ingress and in binding admixed or
intruding chlorides . This was frs t es tablis hed by Verbeck
(1 968) and has s ince been confrmed by other res earch
(Rasheeduzzafar et al. 1 992). The main conclusion of this
work is that the use of very low C 3 A (Type V) cements in a
strong chloride environment is generally not recommended.
9.4.2 Supplementary cementitious materials —Fly ash, slag
cement, and silica fume are generally assumed to improve the
resistance of concrete to chloride-induced corrosion. While
the introduction of such materials to concrete will consume
some of the Ca(OH) 2 that acts as a buffer against changes in
pH due to carbonation of concrete (Bijen and van Selst 1 991 ;
Horiguchi et al. 1 994; Branca et al. 1 992), improvements in
pore distribution and permeability can counteract this depletion in CaOH 2 (Torii and Kawamura 1 994; Hakkinen 1 992).
Also, SCMs can increase the electrical resistivity of the
uncracked concrete, thus reducing the rate of any corrosion
that has been initiated (Schiessl et al. 1 994; Fidjestøl 1 987,
1 991 ; Alonso et al. 1 992).
50
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
In the 1 990s, the development of corrosion-resistant
concrete focused on using blends of portland cement and
other cementitious materials (Baweja et al. 1 994; Maage and
Helland 1 991 ; Maage et al. 1 994; Berke et al. 1 991 ; Collepardi
et al. 1 994; Anqi et al. 1 991 ; Hussain and Rasheeduzzafar
1 994; Decter et al. 1 989; Haque et al. 1 992; Ozyildirim 1 994).
Malhotra et al. (2000) and Smith et al. (2004) all reported
that concrete containing moderate to high volumes o f f y as h
exhibited superior resistance to the penetration of chlorides
and improved corrosion resistance.
Slag cement has been used in marine work since the early
1 900s; the experiences with respect to resistance against
chloride-induced corrosion are generally good (Wiebenga
1 984; Hope et al. 1 985 ; Pal et al. 2002). Slag cement has also
been shown to improve resistance to penetration of deicer
salts (McGrath and Hooton 1 997; Bleszynski et al. 2002).
Silica fume works in several ways to reduce the risk of
corrosion. The reduced permeability of silica fume concrete
means a greatly reduced rate of chloride penetration in
marine structures and structures exposed to deicing salts.
Such concrete also has very high electrical resistivity, thereby
greatly diminishing the rate of corrosion once it is initiated
(Wolsiefer 1 991 ; Pettersson 1 995; Fidjestøl 1 987, 1 993 ;
Fidjestøl and Frearson 1 994; Alonso et al. 1 992; Berke et al.
1 992; Gautefall and Vennesland 1 985; Zhang and Gjørv 1 991 ;
Fidjestøl and Justnes 2002; Skjølsvold et al. 2007).
9.4.3 Aggregates —Aggregates can contain chloride
salts, particularly those aggregates that have been exposed
moresites
FREE
from
to seawater or whoseGet
natural
are standards
in groundwater
containing chloride. Sedimentary rock formed in ancient
s eabeds can also contain s ignifcant amounts o f chlorides .
There have been reported instances (Gaynor 1 985 ) where
quarried stone, gravel, and natural sand contained small
amounts of chloride that have resulted in concrete chloride
contents that exceed the maxima described in 9.2.2.5. Note,
however, that tightly-bound chlorides in aggregate may not
contribute to corrosion of steel. ASTM C1 524 can be used
to determine the water-extractable chloride content of aggregates that could potentially contribute to corrosion initiation.
9.4.4 Mixing water—Potable mixing water can contain
small amounts of chloride. ASTM C1 602/C1 602M has
optional limits for chlorides in mixing water: 500 ppm for
prestressed concrete, bridge decks, or as otherwise designated; and 1 000 ppm for other reinforced concrete in moist
environments.
9.4.5 Admixtures
9.4.5.1 General —Admixtures containing s ignifcant
concentrations of CaCl 2 should not be used in concrete
containing embedded metal. Some water-reducing admixtures can contain chloride to improve admixture performance, but contribute only small amounts of chloride to
the concrete when they are added at recommended rates.
Normal-setting admixtures that contribute much less than
0.1 percent chloride by mass of cement are most common;
their use should be evaluated based on the application.
Chemical admixtures are described in detail in ACI 21 2.3R.
9.4.5.2 Accelerators —Accelerating admixtures, other
than those based on CaCl 2 , have been used in concrete with
varying success. Accelerators that do not contain chloride
should not automatically be assumed to be noncorrosive.
The materials most commonly used in chloride-free accelerators are calcium formate, sodium thiocyanate, calcium
nitrate, and calcium nitrite. It is generally accepted that
formates (Holm 1 987) are noncorrosive in concrete, and that
calcium nitrite is also an inhibitor.
9.4.5.3 Inhibitors —ACI 222.3R provides an overview of
corrosion inhibitors for concrete systems. Four corrosioninhibiting admixtures are common commercially: 1 ) amine
carboxylate; 2) amine-ester organic emulsion; 3) calcium
nitrite; and 4) an organic alkenyl dicarboxylic acid salt (ACI
21 2.3R). Amine carboxylate admixture was developed from
vapor phase inhibitors that have a long history of use in
other industries. As an anodic and cathodic inhibitor, it can
be useful in both new construction and repair applications
(Bavarian and Reiner 2004). Amine-ester organic emulsion
is reported to protect by reducing chloride ingress and by
forming a protective flm at the s teel s urface ( Nmai et al.
1 992; Bobrowski and Youn 1 993 ). Laboratory evaluations
indicate that amine-ester organic emulsion will delay the
onset and reduce the rate of corrosion (Nmai et al. 1 992;
Nmai and Krauss 1 994).
Calcium nitrite has been widely used as an accelerating
admixture that will also function as a corrosion inhibitor.
Laboratory studies have demonstrated that it delays the onset
of corrosion or reduces the rate after it has been initiated
(Berke 1 985 ; Berke and Roberts 1 989). The ratio of chloStandard
our chats
ride ionsSharing
to nitriteGroup
ions is and
important.
Studies (Berke 1 987)
show that calcium nitrite can provide corrosion protection
even at chloride-nitrite ratios exceeding 1 .5 to 1 .0 by mass.
Although dosage rates vary, 2 to 6 gal./yd 3 (1 0 to 30 L/m 3 )
of concrete is the common range. Berke and Rosenberg
(1 989) compiled an extensive review of the use of calcium
nitrite as a corrosion inhibitor for steel, galvanized steel, and
aluminum in concrete, which was later updated by Berke et
al. (1 994). If the accelerating effect from calcium nitrite is
undesirable, use of a retarder is recommended. Increased
amounts of air-entraining admixture may be necessary when
calcium nitrite is used to maintain the desired air content.
Montes et al. (2004) showed that the effect of calcium nitrite
on corrosion inhibition in cracked elements was limited.
9.5—Mitigating corrosion
9.5.1 General —In mitigating corrosion, these critical
points should be evaluated in terms of concrete service life:
a) Initiation of corrosion
b) Corrosion products becoming visible (staining)
c) Serviceability, including cracking, spalling, or both
d) The load-carrying capacity of the structure if it is seriously reduced, or the structure can no longer perform its
intended purpose
Traditional service life considerations only consider the
initiation of corrosion; however, even for very conspicuous
and visible structures, criterion b) or c) may be more appropriate in design.
The corrosion of steel in concrete is dependent on the environment in which the structure is exposed. Table 1 9.3.1 .1
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
in ACI 31 8-1 4 s pecifes expos ure categories for corros ion
protection of reinforcement: C0 is designated for concrete
dry or protected from moisture; C1 is for concrete exposed
to moisture but not to an external source of chlorides; and
C2 is designated for concrete exposed to moisture and an
external source of chlorides from deicing chemicals, salt,
brackish water, seawater, or spray from these sources. Corrosion can be minimized by selecting processes and materials
that delay the onset of corrosion and then minimizing the
rate. Increasingly, service life prediction models are being
used to determine combinations of concreting materials
that can help achieve design service lives from a corrosion
perspective. Information on service life models are found in
ACI 365.1 R. Detailed guidance on preventive strategies can
be found in ACI 222.3R. A general discussion of some of
the factors affecting corrosion resistance are described in the
following sections.
9.5.2 Design and process
9.5.2.1 Concrete quality and cover over steel
9.5.2.1.1 Cover depth —Extensive studies (Clear 1 976;
Pfeifer et al. 1 987; Marusin and Pfeifer 1 985 ) have shown
that 1 in. (25 mm) cover over bare steel bars is inadequate
for chloride protection in severe corrosion environments,
even if the concrete has a w/cm as low as 0.30. Tests have also
shown that the chloride content in the top 1 /2 in. (1 2 mm) of
concrete can be very high compared with that at depths of 1
to 2 in. (25 to 50 mm), even in concrete of high quality such
as one having a w/cm of 0.30. As a result, cover for moderateto-severe corrosion environments should be a minimum of
1 -1 /2 to 2 in. (40 to 50 mm).
Concrete will absorb salts applied in deicing operations. To
postpone initiation of corrosion, cover should be maximized.
Trejo and Reinschmidt (2007) reported that increasing the
concrete cover is the most effective way to increase the time
to corrosion and the service life of a concrete structure. Too
much cover, however, can increase cracking. Weyers et al.
(2003) reviewed the inf uence o f cover depths and recom mended cover depths for bridges of 2.75 in. (70 mm). ACI
31 8-1 4 s pecifes concrete cover requirements for Expos ure
Condition C2 and these depend on additional exposure
conditions, member type, and reinforcement type and size.
9.5.2.1.2 Concrete quality —Numerous test programs have
shown that concrete made with a low w/cm and adequate
cover
over
the
s teel
per forms
s ignifcantly
better
than
concrete made with a higher w/cm . Chloride ion penetration
to a 1 in. (25 mm) depth is approximately 400 to 600 percent
greater for concrete made with w/cm of 0.40 and 0.50 than
for concrete with w/cm of 0.32 (Pfeifer et al. 1 987). Similarly, the proper use of SCMs can extend the time to corrosion and reduce its rate.
ACI 3 1 8 and other s pecifcations
place strict requirements
on the mixture proportions for severe chloride environments.
The Norwegian Public Roads Administration (2009) has
specifc
requirements
for mixture proportions , depending
on the location of the structure (w/cm less than 0.38 to 0.40
and a certain silica fume content) . The s pecifcations for the
Great Belt (Storbælt) project required a low w/cm with the
use o f s ilica fume and f y as h to provide a 1 0 0 - year s ervice
51
Table 9.5.2.1.2—Limits to chlorides in newly
constructed concrete (as recommended by ACI
222R)
*
Acid-soluble
Water-soluble
(ASTM C1152/
(ASTM C1218/
Soxhlet
C1152M)
C1218M)
method *
Prestressed
concrete
0.08
0.06
0.06
Reinforced
concrete in wet
conditions
0.1 0
0.08
0.08
Reinforced
concrete in dry
or protected
conditions
0.20
0.1 5
0.1 5
Soxhlet method described in ACI 222.1 R.
Note: All chloride contents expressed as percent Cl – by mass of cement.
life (Storebælt Technical Publications 1 999). A similar
mixture proportioning philosophy was used for 1 .3 to 2.6
million yd3 (1 to 2 million m3 ) of concrete for the 6 mile (1 0
km) connection across Øresund from Sweden to Denmark
(Henriksen et al. 2000).
Admixed chlorides can als o in
f
uence the quality and long-
term performance of concrete structures. Table 9.5.2.1 .2
gives the limits to chlorides in newly constructed concrete
recommended by ACI 222R.
Note that in s pecifcations in other countries, such as CSA
A23.1 -1 4/CSA A23.2 in Canada and EN206 in Europe,
chloride limits are expressed as percent by mass of total
cementing materials, which is portland cement plus SCMs.
Note that the chloride thresholds for prestressed concrete
are lower than those for reinforced concrete. Prestressing
steels have different chemical composition than conventional reinforcing steel, and often suffer from additional
corrosion mechanisms (9.6). The effect of corrosion mechanisms and high stresses in these steels is to make extra restrictions necessary compared to black steel. Structure type and
importance, fabrication methods, and exposure conditions
als o in f uence corros ion per formance o f thes e s teels .
For post-tensioned structures, the Federal Highway
Administration (FHWA) (201 2) reported that although codes
allow a maximum of 0.08 percent total chloride content
by weight of cement in fresh grout, a lower total chloride
threshold value may be required.
9.5.2.1.3 Cracks —Cracks permit much faster chloride
infltration rate than di ffus ion proces s es , and can es tablis h
chloride concentration cells that accelerate corrosion. ACI
31 8-99 concluded that the role of cracks in the corrosion of
rein forcement is controvers ial and, without s ound s cientifc
data relating
crack width to corrosion
activity,
modifed the
design procedure for limiting crack widths, eliminating the
crack width wmax from the design equation and substituting
an equation based on controlling the bar spacing. Although
ACI 31 8 does not provide explicit limits for crack widths in
the equation, this equation does indirectly limit the maximum
crack width to between 0.01 6 and 0.02 in. (0.4 and 0.52 mm).
In addition, the presence of cracking effects on corrosion
52
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
may be more than just the ingress of deleterious materials, as
when the crack tip reaches the steel; a small anode to large
cathode ratio is created, which may greatly accelerate corrosion, as well as further cracking. In structures submerged in
seawater, for example, large, dynamic (working) cracks have
been reported to be closed by brucite and aragonite (magnesium hydroxide and calcium carbonate) within a fairly short
period (Espelid and Fidjestøl 1 986; Buenfeld and Newman
1 986). ACI 224R and ACI 224.1 R provide good overviews
of issues associated with cracks and corrosion.
To minimize crack formation, concrete should always be
made with the lowest practical water content. Also, proper
detailing o f rein forcement, suffcient minimum and struc tural reinforcement, and control of heat of hydration and
restraint effects are important in producing a structure in
which cracks do not degrade corrosion resistance. Use of
shrinkage-reducing admixtures (SRAs), saturated prewetted
lightweight aggregate for internal curing, and mixture modifcation to reduce paste volume mixture optimization are
additional strategies to reduce cracking in concrete mixtures.
9.5.2.2 Concrete resistivity —When concrete is kept
moderately dry, corrosion of steel is minimized. For
example, if concrete containing up to 2 percent CaCl 2 is
allowed to dry to a maximum internal relative humidity of
50 to 60 percent, embedded steel should either not corrode
or corrode at an inconsequential rate (Tutti 1 982). Maintaining an internal relative humidity below 50 percent,
however, is not always possible.
more
from
While the surface Get
regions
of FREE
exposedstandards
concrete structures will have high or low electrical conductivity values
depending on the wetting and drying conditions of the environment, the concrete interior usually requires long drying
periods to achieve low electrical conductivity. Pfeifer et
al. (1 987) found that 7 to 9 in. (1 80 to 230 mm) thick reinforced concrete slabs with w/cm ranging from 0.30 to 0.50
have essentially equal initial AC electrical-resistance values
between the top and bottom reinforcing bar mats at 28 days.
Cementitious materials that include SCMs can give very
high electrical resistivity in concrete. Slag cement, f y ash,
and, in particular, silica fume, will give concrete resistivities far in excess of what is provided by portland-cement
concrete (Cabrera and Ghoddoussi 1 994; Fidjestøl and
Frearson 1 994; Gautefall and Vennesland 1 985 ; Berke
1 988). Similarly, AC resistance tests on concrete made with
silica fume at a w/cm of 0.20 show extremely high initial
electrical resistance when compared with concrete having a
w/cm of 0.30 to 0.50. The high electrical resistance of silica
fume concrete can be due to denser paste microstructure; to
changes in the pore chemistry; and at low w/cm , to self-deiccation. Field investigations after more than 20 years confrm
long-term performance of SCMs (silica fume) (Fidjestøl and
Justnes 2002; Skjølsvold et al. 2007).
The high electrical resistivity of blended binder systems is
confrmed by tests using ASTM C1 202 (or AASHTO T277),
which provides a method to determine conductivity that is
then used as an indirect indication of chloride diffusivity. In
several investigations, there has been a relationship to chloride diffusion determined by more conventional diffusion
or ponding tests (Detwiler and Fapohunda 1 993 ; Wolsiefer
1 991 ; Misra et al. 1 994; Burg and Öst 1 992). As implied in
the standard, however, this relationship cannot be assumed
to be universal because it also depends on the composition
of the binder system, such as content of and type of cement
and cementitious materials used.
9.5.3 Construction aspects
9.5.3.1 Workmanship —Good workmanship is vital for
securing uniform concrete with low penetrability. For lowslump concrete, segregation and honeycombing can be
avoided by good consolidation. Meeting the requirements of
the specifcations pertaining to durability are essential.
9.5.3.2 Rein forcement detailing —Two factors are important to consider in detailing of the reinforcement:
1 . Adequate spacing should be provided to allow for
proper placing of the concrete cover so that honeycombing
and poor compaction are avoided and good bond between
concrete and steel are obtained.
2. Corrosion is relatively more severe for small bars than
for large bars. Corrosion of a No. 3 (1 0 mm) bar totaling
0.04 in. (1 mm) of corrosion means nearly 40 percent loss of
cross section, whereas for a No. 8 (25 mm) bar, it will mean
1 5 percent loss of cross section. Note, however, that large
bars could cause larger cracks than smaller bars because
smaller bars can give better crack distribution.
9.5.3.3 Curing —Good curing reduces permeability
because of increased hydration of the cement. At least 7 days
of uninterrupted moist curing or membrane curing is ideally
Standard
Sharing
and our
chatsis also important.
specifed.
LimitingGroup
early thermal
stresses
Good curing reduces transport parameters as well as charge
passed in ASTM C1 202 (or AASHTO T277) (Acker et al.
1 986; Marusin 1 989 ). In addition, due to insu ffcient curing,
the chloride penetration rate in the near-surface part of the
concrete cover can be many times greater than at depth
(Hooton et al. 2002).
9.5.3.4 Formwork —Good, tight formwork is essential. Properly supported screeding equipment and correct
supports for the reinforcement are important for attaining the
cover protection specifed. The use o f side form spacers for
reinforcing bars in vertical formwork is similarly important.
Controlled permeability liners for formwork may improve
the quality of the cover (Sha’at et al. 1 993 ).
9.5.4 Design —Design can do much to reduce corrosion
attack because proper detailing can minimize accumulation
of salts and the establishment of high humidity areas where
corrosion can be sustained.
9.5.4.1 General layout o f structure —Placement and
general layout of the structure are important for a favorable
environment. An increase in the height of a bridge over the
sea will reduce the chloride exposure: feld inspection o f
concrete bridges (Fluge and Blankvoll 1 995 ) found that an
increase in bridge height above sea level from 26 to 92 ft
(8 to 28 m) reduces the amount of chlorides deposited on
the surface by as much as 85 percent. The chloride exposure
was also up to eight times higher in the lee side of the structure than on the windward side. Similarly, moving bridge
columns away from traffc splash will reduce the chloride
exposure of the concrete.
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
9.5.4.2 Drainage —Particular attention should be given to
design details to ensure that water will drain and not pond
on surfaces. There are a number of details that are important
(Kompen 1 994), such as proper slope and extended drainage
pipes that take the water away from the concrete surface.
9.5.4.3 Exposed items —Attention should be given to
partially embedded or partially exposed items, such as
bolts, that are exposed directly to corrosive environments.
The resistance of these items to the corrosive environment
should be investigated, and the coupling of dissimilar metals
should be avoided. Concrete should be carefully placed
around embedded items so that it is well consolidated and
does not create paths that will permit corrosive solutions to
easily reach the interior of the concrete.
9.5.5 Special protective systems —Costs of repairing
corrosion-induced damage are very high. Many protective
systems have been proposed, and the reader is referred to
ACI 222R, ACI 222.3R, and ACI 51 5.2R for a comprehensive understanding of protective and mitigating options.
Some of these protective systems have been shown to be
effective. Several of these are listed as follows:
a) Overlays and patches of very low w/cm (0.32), latexmodifed concrete overlays ( Clear and Hay 1 973 ; FHWA
1 975 ), concrete containing silica fume, and concrete
containing high-range water-reducing admixtures
b) Epoxy-coated reinforcing steel
c) Corrosion-resistant steels (Rasheeduzzafar et al.
1 992; Trejo and Pillai 2004; Clemeña and Virmani 2004;
Williamson et al. 2003 )
d) Waterproof membranes (Van Til et al. 1 976)
e) Surface protective-barrier systems produced from
silanes, siloxanes, epoxies, polyurethanes, and methacrylates (Van Daveer and Sheret 1 975 )
f) Cathodic protection
9.6—Corrosion o f prestressed steel rein forcement
The mechanisms and risks associated with general surface
corrosion and pitting in prestressed steel reinforcing systems
are comparable to conventional reinforcing steel (9.1 to 9.5),
with added concerns of fracture due to hydrogen embrittlement and stress corrosion cracking. Hydrogen embrittlement
is the result of a loss of ductility in the steel reinforcement
from the local absorption of atomic hydrogen released from
corrosion cells, and contact with water, hydrogen sulfde,
and other sources at the steel surface. Stress corrosion
cracking is similarly a brittle fracture event caused by the
interaction of the tensile stress within the steel reinforcement and corrosive environment produced as described
previously. Certain elements of prestressing steel systems,
such as end anchorages and slab tendons, that are exposed
to deicing solutions or come in contact with other metal
components (for example, ducts), may have added risks due
to reduced concrete cover protection, galvanic action, or
direct exposure to aggressive environments. Unlike general
surface corrosion, these mechanisms produce sudden loss of
prestressing function and permanent damage to the aff icted
element(s), and are di ffcult to locate via inspection or
testing when the structure is in service. Preventive measures
53
during fabrication/construction are needed to reduce the risk
of occurrence. ACI 222.2R provides addition details on the
prestressing systems, deterioration, protection, feld evalua tion, and remediation techniques.
Corrosion of prestressing steels is best prevented when
protective measures start at the fabrication shop and continue
through proper installation and placement in durable sheath,
concrete, or grout materials. Prestressing steel raw material
and fnished anchorage products should be protected from
exposure to corrosive elements such as rain, snow, deicing
chemicals, salt-spray, and water at the fabrication shop,
during transport to the construction, and throughout the
on-site storage process until placement. Ideally, tendons and
wire should be stored indoors in a climate-controlled warehouse, shipped after shrink-wrapping and coverage with
tarps, and stored on site in a climate-controlled environment.
These actions prevent the formation of any corrosion cells
before placement.
Protection must continue once tendons, strands, or wires
are prepared for installation, placed in the formwork, and
prepared for tensioning. Temporary protection of exposed
tendons or strands and anchorage materials at construction
joints and ends is recommended, as the prestressed reinforcement may be exposed to the atmosphere during the
period when the concrete is placed and reaches the required
strength for stressing. Sheaths and ducts must be watertight, with excess water removed. Any tendon tails or extra
strand length should be cut and capped after inspection is
complete and stressing is started. Before and after stressing,
all exposed surfaces should be inspected and cleaned, with
water removed, and anchor cavities grouted shortly thereafter to avoid exposure. Systems that employ grease canisters at end anchorages should be periodically inspected to
confrm that the grease is not leaking and the canisters are
full. Certain specialized evaluation and repair procedures
have been developed to address select prestressing systems
after being placed in-service. For a more detailed discussion
of the corrosion of prestressing strand, refer to ACI 222.2R.
9.7—Degradation o f materials other than steel
9.7.1 Introduction —Nonferrous metals are occasionally
used in concrete. These metals include aluminum, lead,
copper and copper alloys, zinc, cadmium, Monel metal,
stellite (cobalt-chromium-tungsten alloys), silver, and tin.
Galvanized steel and special alloys of steel, such as stainless
steels and chrome-nickel steels, have also been used. Zinc
and cadmium are used as coatings on steel.
Corrosion of nonferrous metals or alloys can result from
various phenomena. The metal may be unstable in highly
alkaline concrete or in the presence of chloride ions. The
former occurs when the concrete is relatively fresh and may
be self-limiting. The latter can initiate corrosion, particularly when the metal is in contact with a dissimilar metal.
When dissimilar metals are in electrical contact (coupled),
a galvanic cell can occur, resulting in corrosion of the more
active metal. More detailed information on corrosion of
nonferrous metals is available (Fintel 1 984; Erlin 2006).
54
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
9.7.2 Aluminum —Corrosion of aluminum embedded in
concrete can crack the concrete. Conditions conducive to
corrosion are created if the concrete contains steel in contact
with the aluminum, chlorides are present in appreciable
concentrations, or the cement is high in alkali content (Woods
1 968; Erlin 2006). When the metals are coupled, increasing
ratios of steel area, particularly in the presence of appreciable
chloride concentrations, increase corrosion of the aluminum.
Additionally, hydrogen gas evolution may occur when fresh
concrete contacts aluminum. This may increase the porosity of
the concrete and, therefore, the penetration of future corrosive
agents. Some aluminum alloys are more susceptible to this
problem than others. Corrosion inhibitors, such as calcium
nitrite, have been shown to improve the corrosion resistance
of aluminum in concrete (Berke and Rosenberg 1 989).
9.7.3 Lead—Lead in damp concrete will be attacked by
the CaOH 2 in the concrete, and may be destroyed in a few
years. Contact of lead with reinforcing steel can accelerate
the attack. Lead should be isolated from the concrete by
protective plastic, or other materials that are unaffected by
damp concrete. Corrosion of embedded lead is not likely to
damage the concrete (Alhassan 2005 ).
9.7.4 Copper and copper alloys —Copper is not normally
corroded by concrete, as is evidenced by the widespread and
successful use of copper waterstops and the embedment of
copper pipes in concrete for many years (Erlin 2006). Corrosion of copper pipes, however, has been reported where
ammonia is present. Also, there have been reports that small
FREE
standards
from
amounts of ammonia, Get
and more
possibly
of nitrates,
can cause
stress corrosion cracking of embedded copper. Galvanic
corrosion of steel will occur if the steel is connected to the
copper (Erlin 2006).
9.7.5 Zinc —Zinc reacts with alkaline materials, such as
those found in concrete. Zinc in the form of a galvanizing
coating on reinforcing steel, however, is sometimes intentionally embedded in concrete. Available
data are con
f
icting
f any, o f this coating ( Cook 1 980; Stark
and Perenchio 1 975 ; Hill et al. 1 976; Gri ffn 1 9 6 9 ; Federal
Highway Administration 1 976). A chromate dip on the
galvanized bars or the use of 400 ppm of chromate in the
mixing water is recommended to prevent hydrogen evolution in the fresh concrete. Use caution when using chromium
salts because of possible skin allergies. Additionally, users
are cautioned against permitting galvanized and black steel
to come in contact with each other in a structure because the
use of dissimilar metals can cause galvanic corrosion. Corrosion inhibitors, such as calcium nitrite, have been shown to
improve the corrosion resistance of zinc in concrete (Berke
and Rosenberg 1 989; Page et al. 1 989).
9.7.6 Other metals —Chromium and nickel alloys generally have good resistance to corrosion in concrete, as do
silver and tin. The corrosion resistance of some of these
metals may be adversely affected by the presence of soluble
chlorides from seawater or deicing salts. Use of stainless
as
to
steel
the
may
beneft,
be
i
economically
j ustifed
in
some
high
chloride
environments where the higher initial cost is offset by reduced
cost in service over the life cycle. Examples would be marine
locations and heavily deiced bridge decks. The 300 Series
stainless steels, however, are susceptible to stress corrosion
cracking when the temperature is over 1 40°F (60°C) and chloride solutions are in contact with the steel material. Embedded
natural-weathering steels generally do not perform well in
concrete containing moisture and chloride. Weathering steels
adjoining concrete may discharge rust and cause staining of
concrete surfaces (McDad et al. 2000).
9.7.7 Polymers —Polymers are being used increasingly
in concrete in applications such as pipes, shields, reinforcement, waterstops, chairs, and concrete reinforcement. Many
plastics are resistant to strong alkalis and would, therefore,
be expected to perform satisfactorily. Because of the great
variety of plastics and materials compounded with them,
however,
s pecifc tes t data s hould be developed for each
intended use.
9.7.8 Wood—Wood has been widely used in or against
mortars and concrete. Such use includes the incorporation of
sawdust, wood pulp, and wood fbers in the concrete and the
embedment of timber (Erlin 2006).
The
us e
o f untreated
sawdust,
wood
chips ,
or
fbers
usually results in slow-setting and low-strength concrete
(Erlin 2006). The addition of hydrated lime equal to onethird to one-half the volume of the cement is usually effective in minimizing these problems. The further addition of
up to 5 percent of CaCl 2 dihydrate by mass of cement has
also helped to minimize these problems (Erlin 2006). CaCl 2
in such amounts can cause corrosion of embedded metals,
however, and can have adverse effects on the concrete.
Standard
Sharing
Group
and concrete
our chats
Another
problem
with such
is the high volume
change, which occurs even with changes in atmospheric
humidity. This volume change may lead to cracking and
warping (Erlin 2006).
The embedment of lumber in concrete has sometimes
resulted in leaching of the wood by Ca(OH) 2 with subsequent
deterioration. Soft woods, preferably with high resin content,
are reported to be most suitable for such use (Erlin 2006).
9.8—Summary
Portland-cement concrete can provide excellent corrosion
protection to embedded steel. When corrosion occurs, the
costs of repairs can be exceedingly high. The use of highquality concrete; adequate cover over the steel; appropriate
reinforcement type; and proper design, including detailing
and additional protection, are prerequisites if deterioration
due to steel corrosion is to be minimized.
ACI 222R and ACI 222.3R provide a summary of the
causes and mechanisms of corrosion of steel. They include
information on how to protect against corrosion in new
structures and procedures for identifying corrosive environments and remedial measures where corrosion is occurring.
CHAPTER 10—ABRASION
10.1 —Introduction
The abras ion res is tance o f concrete is defned as the ability
of a surface to resist being worn away by rubbing and friction.
Abras ion o f f oors and pavements can res ult from production
operations or foot or vehicular tra ffc. Abras ion res is tance
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
is, there fore, o f concern in industrial f oors. Wind, water, or
waterborne particles can also abrade concrete surfaces (Price
1 947). There are instances where abrasion is of little concern
structurally, yet there may be a dusting problem that can be
objectionable in some kinds of service. Abrasion of concrete
in hydraulic structures is discussed only brie f y in this guide;
the subject is treated in more detail in ACI 21 0R and 21 0.1 R.
10.2—Testing concrete for resistance to abrasion
Research to develop meaningful laboratory tests on
concrete abrasion has been ongoing since the early 1 900s.
There are several different types of abrasion, and no single
test method has been found that is adequate for all conditions. A detailed description of abrasion/erosion test methods
can be found in Bakke (2006). Prior (1 966) described four
broad areas related to abrasion:
1 . Wear on f oor and slab construction; Table 1 0.2 shows
classes o f f oor traffc and use, and the special considerations
required for good wear resistance (ACI 302.1 R).
2. Wear on concrete road surfaces due to attrition, scraping,
and percussion from heavy trucks and automobiles.
3. Erosion of hydraulic structures, such as dams, spillways, tunnels, bridge piers, and abutments, due to the action
o f abrasive materials carried by f owing liquid (attrition and
scraping).
4. Cavitation action on concrete in dams, spillways,
tunnels, and other hydraulic structures due to high f ow
velocities and negative pressures.
ASTM C779/C779M covers three operational procedures for evaluating f oor surfaces: Procedure A, revolving
discs; Procedure B, dressing wheels; and Procedure C, ball
bearings. ASTM C944/C944M is similar to ASTM C779/
C779M Procedure B and is used for testing smaller areas
than required for ASTM C779/C779M.
Each method has been used to develop information on
wear resistance. Liu (1 994) commented that the most reproducible results are obtained by the method involving the use
of revolving discs. Reproducibility of abrasion testing is an
important factor in selecting the test method. Replication of
results is necessary to avoid misleading results from single
tests.
The concrete surface condition, loose aggregates that
are dislodged and abraded during the test procedure, and
care and selection of representative samples can all affect
test results. Samples that are fabricated in the laboratory
should be identical for comparison, and the selection of
feld-testing sites should be made on the basis o f providing
representative results.
To set limits for abrasion resistance of concrete, it is necessary to rely on the relative values obtained during testing to
provide a prediction of service.
Underwater abrasion presents special demands for test
procedures used to assess durability. ASTM C11 38 uses agitation of steel balls in water to determine abrasion resistance.
ASTM C41 8 uses a sandblasting apparatus to measure the
depth or wear to simulate comparative sand-impinged wear
resistance. This test provides a means for evaluating resistance to abrasion caused by wind-blown sand.
55
The abrasion resistance of pervious concrete can be
measured using ASTM C1 747/C1 747M . In this test method,
the impact and abrasion resistance of the pervious concrete
is determined from the percent mass loss from three cylindrical concrete specimens after they are placed in a Los
Angeles machine and rotated for 500 revolutions.
The abrasion resistance of railroad crosstie rail seats is
measured using AREMA Test 6 (AREMA 2007). This test
method simulates the abrasion that occurs when moisture
and sand are present between the concrete crosstie rail seat
and pad when heavy loads are applied at an angle. A cyclic
load is applied on the rail at a 27.5-degree angle from vertical
with sand and a water drip placed on each side of each rail
seat. Crossties that have individual components break or a
rail de f ection greater than 0.2 in. (5 mm) during loading
before 1 ,000,000 cycles fail the test (AREMA 2007). This
system test is thought to better simulate the kinetic friction
that causes the pad and grit under the pad to move laterally
back and forth against the crosstie rail seat, causing wear
(Kernes et al. 201 1 ).
In summary, a number of tests previously described aid in
determining the durability of a select concrete mixture design
and sample construction to simulate abrasive actions. Many
o f these tests target specifc exposures or structural types (for
example, railroad ties). Application of one or more of these
tests to assess future performance o f a specifc structural
type to one or more forms of abrasion requires consideration
and alignment of test conditions (for example, sample size
and surface conditions) to expected service conditions.
10.3—Factors affecting abrasion resistance o f
concrete
The abrasion resistance of concrete is a progressive
phenomenon. Initially, the resistance is related to compressive strength of the wearing surface. Therefore, initial judgments regarding relative f oor wear can be made on the basis
of compressive strength.
As so fter paste wears away, however, the particles o f fne
and coarse aggregate are exposed, and abrasion and impact
will cause additional degradation that is more related to the
paste-to-aggregate bond strength and the relative hardness of
the aggregate than to the compressive strength of the concrete.
Tests and feld experience have generally shown that abra sion resistance is proportional to the compressive strength of
concrete (Scripture et al. 1 953 ; Witte and Backstrom 1 951 ).
Because abrasion occurs at the surface, it is critical that the
surface strength be maximized. Resistance can be greatly
improved by the use o f dry shakes and toppings, fnishing
techniques (1 0.4.4), and curing. In addition, the use of
concrete mixtures having low to moderately low w/cm (less
than 0.45) is recommended to improve the strength and wear
resistance of surface paste.
Although useful as a relative indicator, reliance should
not be placed solely on the results of compressive strength
tests. Inspection should be made during installation and
fnishing o f f oor slabs to obtain an abrasion-resistant surface
by encouraging the use o f both power-trowel fnishing and
adequate curing (Kettle and Sadegzadeh 1 987).
56
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
Table 10.2—Floor classifcations* and considerations to improve abrasion resistance
Class
Anticipated traffc type
1 . Exposed
Exposed surface —foot traffc
Use
Special considerations
O ffces, churches, multiunit
residential, decorative
Final fnish
Uni form fnish, nonslip aggregate
in specifc areas, curing
Normal steel-troweled fnish,
nonslip fnish where required
Colored mineral aggregate, color
pigment or exposed aggregate,
stamped or inlaid patterns, artistic
joint layout, curing, surface treatment, maintenance
Burnishing or polishing to
enhance sheen as required
2. Covered
Covered surface —foot traffc
O ffces, churches, commercial,
multiunit residential, institutional with f oor coverings
Flat and level slabs suitably dry for
applied coverings, curing
Light steel-troweled fnish
3. Topping
Exposed or covered
Unbonded or bonded topping
over base slab for commercial
or non-industrial buildings
where construction type or
schedule dictates
Base slab—good uniform level
Base slab—troweled fnish
surface —foot traffc
surface tolerance, curing
Unbonded topping—bondbreaker
on base slab, minimum thickness
3 in. (75 mm), reinforced, curing
Bonded topping—properly
sized aggregate, 3/4 in. (1 9 mm)
minimum thickness curing
4. Institutional/
commercial
Exposed or covered
Exposed surface —industrial
vehicular traffc such as pneumatic wheels and moderately
soft solid wheels
Topping— for exposed surface,
normal steel-troweled fnish;
for covered surface, light
steel-troweled fnish
Institutional or commercial
Level and fat slab suitable for
applied coverings, nonslip aggregate
for specifc areas, curing; coordi nate joints with applied coverings
Normal steel-troweled fnish
Industrial f oors for manu facturing, processing, and
warehousing
Good uniform subgrade, joint
layout, joint load transfer, abrasion
resistance, curing
Hard steel-troweled fnish
surface —foot and light
vehicular traffc
5. Industrial
under unbonded topping;
clean, textured surface under
bonded topping
Industrial f oorsfrom
subjectStandard
to
GoodSharing
uniform subgrade,
joint
Special
metallic or mineral
Get more FREE standards
Group
and our
chats
6. Heavy
industrial
Exposed surface —heavy-
duty industrial vehicular
traffc such as hard wheels and
heavy wheel loads
heavy traffc; can be subject to
impact loads
layout, joint load transfer required,
abrasion resistance, curing
aggregate surface hardener;
repeated hard steel-troweling
7. Heavy
industrial
topping
Exposed surface —heavy-
Bonded two-course f oors
subject to heavy traffc and
impact
Base slab—good uniform
Clean, textured base slab
surface suitable for subsequent bonded topping. Special
power f oats for topping are
optional, hard steel-troweled
fnish
duty industrial vehicular
traffc such as hard wheels and
heavy wheel loads
subgrade, reinforcement, joint
layout, level surface, curing
Topping —composed of well-
graded all-mineral or all-metallic
aggregate. Minimum thickness
3 /4 in. (1 9 mm)
Mineral or metallic aggregate
surface hardener applied to highstrength plain topping to toughen,
curing
*
8.Commercial/
industrial
topping
As in Classes 4, 5, or 6
Unbonded topping—on new or
Bondbreaker on base slab,
minimum thickness 4 in. (1 00
mm), abrasion resistance, curing
As in Classes 4, 5, or 6
9.Critical
surface profle
Exposed surface —super f at
or critical surface tolerance
required; special materialshandling vehicles or robotics
requiring specifc tolerances
Narrow-aisle, high-bay warehouses; television studios, ice
rinks, or gymnasiums (ACI
360R)
Varying concrete quality requirements. Special application procedures and strict attention to detail
are recommended when shake-on
hardeners are used. F F 50 to F F
1 25, super f at f oor, curing
Strictly following techniques as
indicated in 8.9 of ACI 302.1 R
old foors where construction
sequence or schedule dictates
Taken from Table 4.1 of ACI 302.1 R-1 5.
With a given concrete mixture, compressive strength at
the surface is improved by:
a) Avoiding segregation
b) Eliminating bleeding
c) Properly timed fnishing
d) Minimizing surface w/cm (water addition to the surface
during fnishing should not be permitted)
e) Hard troweling of the surface, which should not be done
on concrete containing an air-entraining admixture or having
a total air content greater than 3 percent (refer to ACI 302.1 R)
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
f) Proper curing procedures
Economical proportioning of the mixture for increased
compressive strength includes using a limit on the maximum
w/cm and proper aggregate size. When supplementary
cementitious materials (SCMs) such as silica fume, slag, or
f y ash are used in concrete, the abrasion resistance is generally related to the compressive strength developed (Keshari
2009; Naik et al. 1 995 , 2002; Turk and Karatas 2011 ; Yen
et al. 2007). The use of good curing procedures with SCMs
is essential to ensure adequate abrasion resistance (Yen et
al. 2007). Polymer concrete, polymer-impregnated concrete
(Holland and Gutschow 1 987), epoxy concrete (Mirza et al.
1 990), calcium aluminate cement (Scrivener et al. 1 999),
and calcium sulfoaluminate cement (Markey et al. 2006)
have shown exceptional abrasion resistance. Consideration
should be given to the quality of the aggregate (Scripture
et al. 1 953 ; Smith 1 958). The service life of some concrete
slabs, such as warehouse f oors that are subjected to abrasion
by steel or hard rubber-wheeled traffc, is greatly lengthened
by the use of hard, tough aggregates.
The abrasion resistance of lightweight concrete is a function of the concrete compressive strength; however, the
use of only lightweight aggregates may not be advisable
for structures with high abrasion resistance requirements
(ACI 21 3R). The abrasion resistance of concrete containing
recycled aggregate will depend greatly on the compressive
strength and aggregate type of the recycled concrete (Ekolu
et al. 201 2; de Brito 201 0).
Abrasion-resistant aggregates can be used either with the
dry-shake method (ACI 302.1 R) or as part of a high-strength
topping mixture. If abrasion is the principal concern, addition of high-quality quartz, traprock, or emery aggregates
properly proportioned with cement will increase the abrasion resistance by improving the compressive strength at
the surface. The aggregates used in topping mixtures or dry
shakes should be harder than the aggregate in the concrete.
For additional abrasion resistance, a change to a blend of
metallic aggregate and cement will further increase the abrasion resistance and increase service life. Another advantage
of using metallic aggregate is improved impact resistance,
especially at joints.
The use o f two-course foors using a high-strength topping
is generally limited to foors where both abrasion and impact
resistance are required. While providing excellent abrasion
resistance, a two-course foor will generally be more expen sive. Additional impact resistance can be obtained by using a
topping that contains portland cement and metallic aggregate.
A key element in the production o f satis factory f oor
surfaces is curing (Liu 1 994; ACI 302.1 R; ACI 308R).
Because the uppermost part of the concrete surface is the
region that is abraded by traffc, maximum strength and
toughness are the most important elements for ensuring
resistance to surface abrasion. This is partially accomplished
through proper fnishing operations, troweling techniques,
and adequate and timely curing practices (1 0.4.4). The effect
o f curing e ffciency (absorptivity) at the top-wearing surface
has been shown to be directly related to abrasion resistance.
57
Curing has less effect on the abrasion resistance of deeper
sections of the same concrete (Senbetta and Scholer 1 984).
10.4—Recommendations for obtaining abrasionresistant concrete sur faces
10.4.1 Factors a ffecting abrasion resistance —The
following factors directly impact concrete strength and,
therefore, abrasion resistance (ACI 302.1 R):
a) A low w/cm at the surface— Steps to lower w/cm include
the use of water-reducing admixtures, mixture proportions to
reduce bleeding, timing o f fnishing operations that avoid the
addition of water during troweling, and vacuum dewatering.
b) Well-graded fne and coarse aggregates (meeting
ASTM C33/C33M)—The maximum size of coarse aggregate
should be chosen for optimum workability and minimum
water content.
c) Lowest slump consistent with proper placement and
consolidation as recommended in ACI 309R.
d) Air content consistent with exposure conditions— In
addition to a detrimental effect on compressive strength, air
content levels can contribute to surface blistering and delamination if fnishing operations are improperly timed. Entrained
air should not be used for dry-shake toppings unless special
precautions provided by the manufacturer are followed.
10.4.2 Two-course foors —High-strength toppings in
excess of 6000 psi (40 MPa) provide increased abrasion
resistance. The nominal maximum aggregate size for topping
mixtures is 1 /2 in. (1 2.5 mm).
10.4.3 Special concrete aggregates —Selection of hard
aggregates for improved strength performance at a given w/
cm also improves abrasion resistance. Typically, aggregates
are applied as dry shakes or in high-strength, bonded toppings.
10.4.4 Proper fnishing procedures— Floating and troweling operations should be delayed until the concrete has lost
its surface sheen. It may be necessary to remove free water
from the surface to permit fnishing operations to continue
before the base concrete hardens. Standing water should
never be worked into concrete surfaces because it reduces the
compressive strength of the surface paste. The delay period
will vary greatly depending on temperature, humidity, air
movement, and SCMs used. Greater detail regarding proper
fnishing operations is provided in ACI 302.1 R.
10.4.5 Vacuum dewatering —Vacuum dewatering is a
method for removing water from concrete immediately
after placement (New Zealand Portland Cement Association
1 975 ). While this permits a reduction in w/cm , the quality of
the fnished surface is still highly dependent on the timing o f
fnishing and subsequent actions by the contractor. Ensuring
that proper dewatering is accomplished at the edges of the
vacuum mats is essential. Improperly dewatered areas are
less resistant to abrasion because of a localized higher w/cm .
10.4.6 Special dry shakes and toppings —When severe
wear is anticipated, the use of special dry shakes or topping
mixtures should be used. The manufacturer recommendations should be followed. Additional guidance is provided
in ACI 302.1 R.
10.4.7 Proper curing procedures —For most concrete
f oors, water curing (keeping the concrete continuously wet)
58
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
is the most effective method of producing a hard, dense
surface (Shurpali et al. 201 2). Water curing, however, may
not be always practical. Curing compounds, which reduce
moisture loss in concrete from evaporation, are used as
an alternative. Curing compounds also provide protection
against early carbonation and prevent premature or excessive loss of surface moisture. Moist curing of metallic shake
toppings is not recommended because some water sources
and rainwater have a pH of less than 7, which may result
in the oxidation of the metallic aggregate particles. For
pervious concrete, it was found that curing with a plastic
sheet provided improved abrasion resistance. Application of
soybean oil or a curing compound to the pervious concrete
surface was also found to increase abrasion resistance
(Kevern et al. 2009 ). Latex-modifed concrete was found to
greatly improve the abrasion resistance of pervious concrete
(Wu et al. 2011 ).
Water curing is accomplished through the use of sprays,
ponding, or wet coverings such as damp burlap, or cotton
mats. Water-resistant paper or plastic sheets are satisfactory,
provided that the concrete is frst wetted and then immedi ately covered, with the edges overlapped and sealed using
water-resistant tape. The use of plastic sheeting without a
damp cloth layer can result in nonuniform surface color.
Curing compounds should meet the minimum requirements of ASTM C309 or ASTM C1 31 5 . They should
be applied in a uniform coat immediately after concrete
fnishing and in accordance with the manu facturer’s recom Get more
FREE rates
standards
from
mendations. Recommended
coverage
will vary
depending on the surface texture o f the fnished surface. A
smoother f oor surface will have better moisture retention
properties compared with a textured highway slab. Smaller
peaks and valleys result in a lower evaporation rate and,
therefore, require a lower coverage rate. The compound
should be covered with scuff-resistant paper i f the f oor
is subjected to traffc be fore curing is complete. Curing
compounds should not be required for surfaces that receive
paint or f oor tile unless the curing compound is compatible
with these materials.
Wet curing is recommended for concrete with a low w/cm to
supply additional water for cement hydration, where cooling of
the surface is desired, where concrete will later be bonded,
or where liquid hardeners will be applied. Curing methods
are described in detail in ACI 308R.
Heaters burning fossil fuels or other sources of carbon
dioxide (CO 2 ), such as fnishing machines, vehicles, and
welding machines, should not be used without attention to
proper ventilation. CO 2 can adversely affect fresh concrete
surfaces between the time of placement and application of a curing compound through a mechanism referred
to as carbonation. The severity of the effect is dependent
on the concentration of CO 2 , the humidity and ambient
temperature, and the length of exposure to the air (Kauer
and Freeman 1 955 ; Matsuzawa et al. 201 0). Early carbonation will greatly reduce the abrasion resistance of concrete
surfaces. The extent of the reduction depends on the depth of
carbonation. The only effective repair is to grind the surface
to sound, hard concrete.
10.5—Studded tire and tire chain wear on concrete
Abrasive materials, such as sand, are often applied to the
pavement surface when roads are slippery. Experience from
many years use of sand in winter, however, indicates that
this causes little wear if the concrete is of high quality and
the aggregates are wear-resistant.
Tire chains and studded snow tires, however, can cause
considerable wear to concrete surfaces, even where the
concrete is of high quality. Studded snow tires cause serious
damage, even to high-quality concrete. The damage is due
to the dynamic impact of the small tungsten carbide tip
of the studs, of which there are roughly 1 00 in each tire.
One laboratory study showed that studded tires running
on surfaces to which sand and salt were applied caused
1 00 times as much wear as tires without studs (Krukar and
Cook 1 973 ). Fortunately, the use of studded tires has been
declining for a number of years or is forbidden by law in
certain jurisdictions.
Wear caused by studded tires is usually concentrated in
the wheel tracks. Ruts from 1 /4 to 1 /2 in. (6 to 1 2 mm) deep
can form in a single winter in regions where approximately
30 percent of passenger cars are equipped with studded tires
and traffc is heavy (Smith and Schonfeld 1 970). More severe
wear occurs where vehicles stop, start, or turn (Keyser 1 971 ).
Investigations have been made, principally in Scandanavia, Canada, and the United States, to examine the properties of existing concrete as related to studded tire wear
(Smith and Schonfeld 1 971 ; Keyser 1 971 ; Preus 1 973 ;
Standard
Group and
Wehner Sharing
1 966; Thurmann
1 969).our
In chats
some cases, there was
considerable variability in the data and the conclusions of
the different investigators were not in agreement; however,
most found that a hard coarse aggregate and a high-strength
mortar matrix are benefcial in resisting abrasion.
Another investigation was aimed at developing more
wear-resistant types of concrete overlays (Preus 1 973).
Polymer cement concrete and polymer f y-ash concrete
provide better resistance to wear, although at the sacrifce o f
skid resistance. Steel fber concrete overlays were also tested
and showed reduced wear. Exposed fbers can adversely
affect the tire wear.
Although the reported test results show promise, no
affordable concrete surface has yet been developed that will
provide the same service life, when studded tires are used, as
concrete surfaces exposed to plain rubber tire wear. A report
(Brunette and Lundy 1 996) summarizes available data on
pavement wear and on the performance and winter accident
records for studded tire use.
10.6—Skid resistance o f pavements
The skid resistance of concrete pavements depends on
the surface texture of the concrete. There are two types of
surface texture:
1 ) Macrotexture from surface irregularities that are built in
at the time of construction
2) Microtexture from the type and hardness o f fne aggregate
The microtexture is more important at speeds of less than
approximately 50 mph (80 km/h) (Kummer and Meyer
1 967; Murphy 1 975 ; Wilk 1 978 ). At speeds greater than 50
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
mph (80 km/h), the macrotexture becomes quite important
because it is relied on to help prevent hydroplaning.
The skid resistance of concrete pavement initially depends
on the texture built into the surface layer (Dahir 1 981 ). In time,
rubber-tire tra ffc abrades the surface paste, which removes
the benefcial macrotexture
and exposes the coarse- and fne-
aggregate particles. The rate that the surface paste is removed
and the consequences on the skid resistance of a pavement
depends on the depth and quality of the surface paste and the
rock type (toughness) o f the fne and coarse aggregate.
F ine aggregates containing s ignifcant amounts o f s ilicate
minerals in the larger particle sizes will assist in slowing
down surface wear and maintaining the microtexture necessary for satisfactory skid resistance at slow speed (Fowler
and Rached 201 2; Rado 2009). Certain rock types, however,
polish
under
textured
texture,
rubber- tire
limes tone,
the
more
wear.
dolomite,
rapid
the
These
and
include
serpentine;
polis hing.
Where
very
fne-
the
fner
both
the
the
fne
and coarse aggregate are made of these rock types, there
may be a rapid polishing of the entire pavement surface and
a serious reduction in skid resistance. Where only the coarse
aggregate is of the polishing type, the problem is delayed
until the coarse aggregate is exposed by wear. However, if
the coarse aggregate is, for example, a coarse-grained silica
or vesicular slag, the skid resistance may increase when the
aggregate is exposed (Rado 2009).
Macrotexture is important because it prevents hydroplaning. An example of constructing macrotexture in pavement surfaces is placing grooves in the concrete—either
before hardening (tining) or by sawing after the concrete
has hardened—to provide channels for the escape of water
that is otherwise trapped between the tire and the pavement.
The spaces between grooves have to be especially resistant
to surface abrasion and frost action. A high-quality concrete
that is properly fnis hed and cured has the required durability
and abrasion resistance (Ong and Fwa 2008).
10.7—Erosion
Concrete erosion is the progressive removal of mass
from the concrete surface from chemical attack, abrasion,
or cavitation (ACI 21 0R). Chemical attack erosion occurs
when components of the concrete paste or aggregate are
leached or dis s olved. The rate o f degradation can be signif cantly
enhanced
by
f
owing
liquid
due
to
the
increas ed
rate of material removal and the maintenance of a low pH
near the concrete surface. Erosion by abrasion occurs when
suspended solids in the water impact or grind the surface,
causing material loss. Cavitation occurs when the local pressure in a hydraulic system drops below the liquid vaporization pressure, causing the liquid to vaporize and recondense.
As the liquid rapidly recondenses and the bubble/void phase
collapses, very high pressures are created on the concrete
wall, causing damage and material loss (ACI 21 0R).
This document contains a summary of the degradation
mechanism and concrete material properties that increase
the durability, while a more complete coverage of these
subjects is contained in ACI 21 0R. Erosion by chemical
attack is covered in more detail in Chapter 6 of this guide.
59
1 0. 7. 1
A brasion —Suspended
solids can abrade the
concrete surface uniformly, giving a rather smooth, worn
appearance. Factors that increase abrasion include high
water velocities; large, hard, sharp particles; high total
suspended solids content; long periods of exposure; and
concrete shape. Stilling basins, outlet works, locks, and
tunnel linings are common hydraulic structures that experience abrasion/erosion damage (ACI 21 0R; U.S. Bureau of
Reclamation 1 997). Bridge abutments and other structures
placed in rivers
or other
f
owing
bodies
are
als o
candidates
for abrasion/erosion damage.
No concrete is completely immune to abrasion-related
erosion in hydraulic structures. High-quality pastes with a
dense microstructure are critical to make abrasion-resistant
concrete. Hard, abrasion-resistant concrete aggregates are
necessary. Larger coarse aggregates also help increase the
concrete resistance to abrasion erosion (Liu et al. 2006).
High-performance concrete is more resistant to abrasion
damage because of the higher-quality paste and aggregates. Concrete containing silica fume at low w/cm has
been found to increase the resistance to abrasion/erosion
damage. Polymer concrete has been shown to have excellent
resistance to abrasion/erosion (Klieger and Greening 1 969;
Scrivener et al. 1 999; U.S. Bureau of Reclamation 1 997).
1 0.7.2 Cavitation —The most e ffcient way to prevent cavi tation damage in hydraulic structures is to prevent cavitation.
This is best accomplished by considering cavitation in the
preliminary structural and hydraulic design by changing the
structure’s geometry, reducing surface irregularities, reducing
the f ow rate, or by aeration ( Frizell and Mefford 1 991 ). Joints
should be avoided or minimized when possible because they
can increase turbulence. High-pressure, low-velocity systems
are less likely to experience cavitation. Strict construction
tolerances for concrete smoothness may be necessary to
reduce surface irregularities and localized turbulence.
Cavitation is such a damaging mechanism that no material
can be made cavitation proof. The service life of a structure can be extended by the use of more resistant materials.
Cavitation-resistant concrete has similar requirements as
abrasion-resistant concrete. Concrete mixture parameters
that increase the concrete abrasion erosion resistance apply
equally to resisting cavitation damage (MacDonald 2000).
Latex- modifed
concrete
has
als o
been
s hown
to
increas e
cavitation resistance by increasing paste-aggregate bond
(MacDonald 2000).
For cavitation repair, the cause of cavitation should be
addressed, which may include the addition of aeration or
changing
f
ow
characteristics .
When
thes e
methods
are
impractical, the use of higher-quality, erosion-resistant
repair materials may prolong the service life of the structure. Repairs must follow strict tolerances on smoothness.
Silica fume concrete, epoxy-bonded concrete, or polymer
concrete may be used in the repair. Stainless steel surfaces
may also be used to armor the surface, although damage will
still occur. The use of polymer concrete, an epoxy coating,
or stainless steel also has the advantage of a smooth surface
that reduces turbulence (U.S. Bureau of Reclamation 1 997).
60
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
CHAPTER 11 —SUMMARY
The durability of concrete is one of the characteristics that
make it the most commonly used building material in the
world. This guide describes various factors that can inf uence
the durability of concrete, considering the particular mechanisms of deterioration in the context of the environmental
conditions to which the concrete is to be subjected. Strategies
are presented to increase the durability of concrete through
the use of appropriate materials and mixture proportions,
and emphasizes that appropriate placement practices and
workmanship are also essential to the production of durable
concrete in a given environmental exposure. Specifcally,
this guide discusses the importance of concrete’s resistance
to f uid ingress as it inf uences durability and provides indi vidual chapters on distress mechanisms including freezing
and thawing, alkali-aggregate reaction (AAR), sulfate attack,
aggressive chemical attack, physical salt attack, corrosion of
metals and other embedded materials, as well as abrasion.
For each of these mechanisms of distress, recommendations
are made for preventing or minimizing damage.
CHAPTER 1 2—REFERENCES
ACI committee documents and documents published by
other organizations are listed frst by document number, full
title, and year of publication followed by authored documents listed alphabetically.
Get more
FREE
standards from
ACI 207.1 R-05(1 2)—Guide
to Mass
Concrete
ACI 207.2R-07—Report on Thermal and Volume Change
Effects on Cracking of Mass Concrete
ACI 21 0R-93(08)—Erosion of Concrete in Hydraulic
Structures
ACI 21 0.1 R-94—Compendium of Case Histories
on Repair of Erosion-Damaged Concrete in Hydraulic
Structures
ACI 21 2.3R-1 0—Report on Chemical Admixtures for
Concrete
ACI 21 3R-1 4—Guide for Structural Lightweight-Aggregate Concrete
ACI 21 6.1 -1 4—Code Requirements for Determining Fire
Resistance of Concrete and Masonry Construction Assemblies
ACI 221 .1 R-98(08)—Report on Alkali-Aggregate Reactivity
ACI 222R-01 (1 0)—Protection of Metals in Concrete
against Corrosion
ACI 222.1 R-96—Provisional Standard Test Method
for Water-Soluble Chloride Available for Corrosion of
Embedded Steel in Mortar and Concrete Using the Soxhlet
Extractor
ACI 222.2R-1 4—Report on Corrosion of Prestressing
Steels
ACI 222.3R-1 1 —Guide to Design and Construction Practices to Mitigate Corrosion of Reinforcement in Concrete
Structures
ACI 223R-1 0—Guide for the Use of Shrinkage-Compensating Concrete
ACI 224R-01 (08)—Control of Cracking in Concrete
Structures
ACI 224.1 R-07—Causes, Evaluation, and Repair of
Cracks in Concrete Structures
ACI 301 -1 6—Specifcations for Structural Concrete
ACI 302.1 R-1 5—Guide for Concrete Floor and Slab
Construction
ACI 304R-00(09)—Guide for Measuring, Mixing, Transporting, and Placing Concrete
ACI 305R-1 0—Guide to Hot Weather Concreting
ACI 306R-1 6—Guide to Cold Weather Concreting
ACI 308R-1 6—Guide to External Curing of Concrete
ACI 309R-05—Guide for Consolidation of Concrete
ACI 31 8-99—Building Code Requirements for Structural
Concrete and Commentary
ACI 31 8-1 4—Building Code Requirements for Structural
Concrete and Commentary
ACI 350-06—Code Requirements for Environmental
Engineering Concrete Structures and Commentary
ACI 350.1 -1 0—Specifcation for Tightness Testing o f
Environmental Engineering Concrete Containment Structures and Commentary
ACI 350.2R-04—Concrete Structures for Containment of
Hazardous Materials
ACI 357.1 R-91 (97)—Report on Offshore Concrete Structures for the Arctic
ACI 360R-1 0—Guide to Design of Slabs-on-Ground
ACI 365.1 R-00—Service-Life Prediction
ACI 51 5.2R-1 3—Guide to Selecting Protective Treatments for Concrete
Standard Sharing Group and our chats
ASTM International
ASTM C33/C33M-1 6—Standard Specifcation for
Concrete Aggregates
ASTM C1 09/C1 09M-1 6—Standard Test Method for
Compressive Strength of Hydraulic Cement Mortars (Using
2-in. or [50-mm] Cube Specimens)
ASTM C11 4-1 5—Standard Test Methods for Chemical
Analysis of Hydraulic Cement
ASTM C1 50/C1 50M-1 6—Standard Specifcation for
Portland Cement
ASTM C227-1 0—Standard Test Method for Potential
Alkali Reactivity of Cement-Aggregate Combinations
(Mortar Bar Method)
ASTM C260/C260M-1 0a(201 6)—Standard Specifcation
for Air-Entraining Admixtures for Concrete
ASTM C265-08—Standard Test Method for WaterExtractable Sulfate in Hydrated Hydraulic Cement Mortar
ASTM C289-07—Standard Test Method for Potential
Alkali-Silica Reactivity of Aggregates (Chemical Method)
(withdrawn 201 6)
ASTM C295/C295M-1 2—Standard Guide for Petrographic Examination of Aggregates for Concrete
ASTM C309-11 —Standard Specifcation for Liquid
Membrane-Forming Compounds for Curing Concrete
ASTM C41 8-1 2—Standard Test Method for Abrasion
Resistance of Concrete by Sandblasting
ASTM C441 /C441 M-11 —Standard Test Method for
Effectiveness of Pozzolans or Ground Blast-Furnace Slag
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
in Preventing Excessive Expansion of Concrete Due to the
Alkali-Silica Reaction
ASTM C452/C452M-1 5—Standard Test Method for
Potential Expansion of Portland-Cement Mortars Exposed
to Sulfate
ASTM C457/C457M-1 2—Standard Test Method for
Microscopical Determination of Parameters of the Air-Void
System in Hardened Concrete
ASTM C494/C494M-1 6—Standard Specifcation for
Chemical Admixtures for Concrete
ASTM C563-1 6—Standard Test Method for Approximation of Optimum SO 3 in Hydraulic Cement Using Compressive Strength
ASTM C586-11 —Standard Test Method for Potential
Alkali Reactivity of Carbonate Rocks for Concrete Aggregates (Rock-Cylinder Method)
ASTM C595/C595M-1 5—Standard Specifcation for
Blended Hydraulic Cements
ASTM C61 8-1 5—Standard Specifcation for Coal Fly Ash
and Raw or Calcined Natural Pozzolan for Use in Concrete
ASTM C642-1 3—Standard Test Method for Density,
Absorption, and Voids in Hardened Concrete
ASTM C666/C666M-1 5—Standard Test Method for
Resistance of Concrete to Rapid Freezing and Thawing
ASTM C672/C672M-1 2—Standard Test Method for
Scaling Resistance of Concrete Surfaces Exposed to Deicing
Chemicals
ASTM C779/C779M-1 2—Standard Test Method for
Abrasion Resistance of Horizontal Concrete Surfaces
ASTM C845/C845M-1 2—Standard Specifcation for
Expansive Hydraulic Cement
ASTM C856-1 4—Standard Practice for Petrographic
Examination of Hardened Concrete
ASTM C944/C944M-1 2—Standard Test Method for
Abrasion Resistance of Concrete or Mortar Surfaces by the
Rotating-Cutter Method
ASTM C989/C989M-1 4—Standard Specifcation for
Slag Cement for Use in Concrete and Mortars
ASTM C1 01 2/C1 01 2M-1 5—Standard Test Method for
Length Change of Hydraulic-Cement Mortars Exposed to a
Sulfate Solution
ASTM C1 01 7/C1 01 7M-1 3—Standard Specifcation for
Chemical Admixtures for Use in Producing Flowing Concrete
ASTM C1 038/C1 038M-1 4—Standard Test Method for
Expansion of Hydraulic Cement Mortar Bars Stored in Water
ASTM C11 05-08—Standard Test Method for Length
Change of Concrete Due to Alkali-Carbonate Rock Reaction
ASTM C11 38-97—Standard Test Method for Abrasion
Resistance of Concrete (Underwater Method)
ASTM C1 1 52/C11 52M-04(201 2)—Standard Test Method
for Acid-Soluble Chloride in Mortar and Concrete
ASTM C1 1 57/C1 1 57M-1 1 —Standard Performance Specifcation for Hydraulic Cement
ASTM C1 202-1 2—Standard Test Method for Electrical
Indication of Concrete’s Ability to Resist Chloride Ion
Penetration
ASTM C1 21 8/C1 21 8M-1 5—Standard Test Method for
Water-Soluble Chloride in Mortar and Concrete
61
ASTM C1 240-1 5—Standard Specifcation for Silica
Fume Used in Cementitious Mixtures
ASTM C1 260-1 4—Standard Test Method for Potential
Alkali Reactivity of Aggregates (Mortar-Bar Method)
ASTM C1 293-08(201 5)—Standard Test Method for
Determination of Length Change of Concrete Due to AlkaliSilica Reaction
ASTM C1 31 5-11 —Standard Specifcation for Liquid
Membrane-Forming Compounds Having Special Properties
for Curing and Sealing Concrete
ASTM C1 524-02(201 0)—Standard Test Method for
Water-Extractable Chloride in Aggregate (Soxhlet Method)
ASTM C1 543-1 0—Standard Test Method for Determining
the Penetration of Chloride Ion into Concrete by Ponding
ASTM C1 556-1 1 —Standard Test Method for Determining the Apparent Chloride Di ffusion Coe ffcient o f
Cementitious Mixtures by Bulk Diffusion
ASTM C1 567-1 3—Standard Test Method for Determining the Potential Alkali-Silica Reactivity of Combinations of Cementitious Materials and Aggregate (Accelerated
Mortar-Bar Method)
ASTM C1 580-1 5—Standard Test Method for WaterSoluble Sulfate in Soil
ASTM C1 585-1 3—Standard Test Method for Measurement of Rate of Absorption of Water by Hydraulic-Cement
Concretes
ASTM C1 602/C1 602M-1 2—Standard Specifcation for
Mixing Water Used in the Production of Hydraulic Cement
Concrete
ASTM C1 747/C1 747M-1 3—Standard Test Method for
Determining Potential Resistance to Degradation of Pervious
Concrete by Impact and Abrasion
ASTM C1 778-1 6—Standard Guide for Reducing the Risk
of Deleterious Alkali-Aggregate Reaction in Concrete
American Association o f State and Highway Transportation
Offcials (AASHTO)
PP065-1 1 —Standard Practice for Determining the Reactivity of Concrete Aggregates and Selecting Appropriate
Measures for Preventing Deleterious Expansion in New
Concrete Construction
T259-02—Standard Method of Test for Resistance of
Concrete to Chloride Ion Penetration
T277-1 5—Standard Method of Test for Electrical Indication of Concrete’s Ability to Resist Chloride Ion Penetration
CSA Group
CAN/CSA A3000-1 3—Cementitious Materials Compendium
CSA A23.1 -1 4/CSA A23.2-1 4—Concrete Materials and
Methods of Concrete Construction/Test Methods and Standard Practices for Concrete
CAN/CSA
CSA A23.2-1 4A-1 4—Potential Expansivity of Aggregates; Procedure for Length Change Due to Alkali-Aggregate Reaction in Concrete Prisms
CSA A23.2-26A-1 4—Determination of Potential AlkaliCarbonate Reactivity of Quarried Carbonate Rocks by
Chemical Composition
62
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
CSA A23.2-27A-1 4—Standard Practice to Identify Potential for Alkali-Reactivity of Aggregates and Measures to
Avoid Deleterious Expansion in Concrete
CSA A23.2-28A-1 4—Standard Practice for Laboratory
Testing to Demonstrate the Effectiveness of Supplementary
Cementing Materials and Lithium-Based Admixtures to
Prevent Alkali-Silica Reaction in Concrete
German Institute o f Standardization
DIN 4030:2006-06—Assessment of Water, Soil, and
Gases for Their Aggressiveness to Concrete – Part 1 : Principles and Limiting Values
U.S. Army Corps o f Engineers
CRD-C 662:201 0—Determining the Potential AlkaliSilica Reactivity of Combinations of Cementitious Materials, Lithium Nitrate Admixture and Aggregate (Accelerated Mortar-Bar Method)
EM 1 1 1 0-2-2000:1 994—Standard Practice for Concrete
for Civil Works Structures
Authored documents
American Association of State Highway and Transportation O ffcials, 2000, “AASHTO Guide Specifcation for
Highway Construction, Section 56X, Portland Cement
Concrete Resistant to Excessive Expansion Caused by
Alkali-Silica Reaction,” http://leadstates.tamu.edu/ASR/
library/gspec.stm
Get
FREE standards
from
Abbas, A.; Carcasses,
M.;more
and Ollivier,
J.-P., 1 999, “Gas
Permeability of Concrete in Relation to its Degree of Saturation,” Materials and Structures , V. 32, No. 1 , Jan.-Feb., pp.
3-8. doi: 1 0.1 007/BF02480405
Acker, P.; Foucrier, C.; and Malier, Y., 1 986, “Temperature-Related Mechanical Effects in Concrete Elements and
Optimization of the Manufacturing Process,” Concrete at
Early Ages , SP-95, J. F. Young, ed., American Concrete
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Alhassan, S. J., 2005, “Corrosion of Lead and Lead
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Alksnis, F. F., and Alksne, V. I., 1 986, “Sur le role de la
phase siliceuse dans les processus de destruction de le pierre
de la cement dans les milieux de sulfate,” Proceedings, 8th
International Congress on Chemistry o f Cement, Rio de
Janeiro, Brazil, pp. 1 70-1 74.
Almond, D. K., and Janssen, D. J., 1 991 , “The Washington
Hydraulic Fracture Test for Concrete Aggregates Exposed
to Freezing and Thawing,” Supplemental Papers , Second
CANMET/ACI International Conference on Durability of
Concrete, Montreal, QC, Canada, pp. 265-293.
Alonso, C.; Andrade, C.; Bacle, B.; and Fidjestøl, P.,
1 992, “Corrosion de Armaduras en Microhormigones de
Humo de Silica Carbonatados (Corrosion of Reinforcement
in Carbonated Microconcrete with Silica Fume),” Proceedings, ERMCO Conference, Madrid. (in Spanish)
Andersson, K.; Allard, B.; Bengtsson, M.; and Magnusson,
B., 1 989, “Chemical Composition of Cement Pore Solu-
tions,” Cement and Concrete Research , V. 1 9, No. 3, pp.
327-332. doi: 1 0.1 01 6/0008-8846(89)90022-7
Andrade, C., 1 993, “Calculation of Chloride Diffusion
Coe ffcients in Concrete from Ionic Migration Measure ment,” Cement and Concrete Research , V. 23, No. 3, pp.
724-742. doi: 1 0.1 01 6/0008-8846(93)90023-3
Andrade, C., and Whiting, D., 1 996, “A Comparison of
Chloride Ion Di ffusion Coe ffcients Derived from Concen tration Gradients and Non-Steady State Accelerated Ionic
Migration,” Materials and Structures , V. 29, No. 8, pp.
476-484. doi: 1 0.1 007/BF02486282
Ann, K. Y.; Jung, H. S.; Kim, H. S.; Kim, S. S.; and Moon,
H. Y., 2006, “Effect of Calcium Nitrite-Based Corrosion
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530-535. doi: 1 0.1 01 6/j.cemconres.2005.09.003
Anqi, L.; Baoyu, L.; Guoping, H.; Yeibo, C.; and Guolian,
S., 1 991 , “Study on Corrosion Prevention in Reinforced
Concrete Containing Condensed Silica Fume and Its Application,” Durability o f Concrete , Proceedings of the Second
CANMET/ACI International Conference, SP-1 26, V. M.
Malhotra, ed., American Concrete Institute, Farmington
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AREMA, 2007, “Ties,” Chapter 30, Manual for Railway
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Asgeirsson, H., and Gudmundsson, G., 1 979, “Pozzolanic
Activity of Silica Dust,” Cement and Concrete Research , V.
Standard
Group
our
chats
9, No. 2,Sharing
pp. 249-252.
doi: and
1 0.1 01
6/0008-8846(79)90031
-0
Attiogbe, E. K.; Nmai, C. K.; and Gay, F. T., 1 992, “AirVoid System Parameters and Freeze-Thaw Durability of
Concrete Containing Superplasticizers,” Concrete Interna tional, V. 1 4, No. 7, July, pp. 57-61 .
Backstrom, J. E.; Burrows, R. W.; Mielenz, R. C.; and
Wolkodoff, V. E., 1 958a, “Origin, Evolution, and Effects
o f the Air Void System in Concrete, Part 2—In f uence o f
Type and Amount of Air-Entraining Agent,” ACI Journal
Proceedings , V. 55, No. 8, Aug., pp. 261 -272.
Backstrom, J. E.; Burrows, R. W.; Mielenz, R. C.; and
Wolkodoff, V. E., 1 958b, “Origin, Evolution, and Effects of
the Air Void System in Concrete, Part 3—In f uence o f WaterCement Ratio and Compaction,” ACI Journal Proceedings ,
V. 55, No. 8, Aug., pp. 359-375.
Backstrom, J. E.; Burrows, R. W.; Wolkodoff, V. E.;
and Powers, T. C., 1 954, discussion of “Void Spacing as a
Basis for Producing Air-Entrained Concrete,” ACI Journal
Proceedings , V. 51 , No. 4, Dec., pp. 760-761 .
Baer, J., 1 988, Dynamics o f Fluids in Porous Media ,
Dover, Mineola, NY, 764 pp.
Bakharev, T.; Sanjayan, J. G.; and Cheng, Y. B., 2003,
“Resistance of Alkali-Activated Slag Concrete to Acid
Attack,” Cement and Concrete Research , V. 33, No. 1 0, Jan.,
pp. 1 607-1 61 1 . doi: 1 0.1 01 6/S0008-8846(03)001 25-X
Bakke, K. J., 2006, “Abrasion Resistance,” Signifcance
o f Tests and Properties o f Concrete and Concrete-Making
Materials ,” ASTM STP1 69D, ASTM International, West
Conshohocken, PA, pp. 1 84-1 93.
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
Bakker, R., 1 980, “On the Cause of Increased Resistance
of Concrete Made from Blast Furnace Cement to the AlkaliSilica Reaction and to Sulfate Corrosion,” thesis, RWTH,
Aachen, Germany, 1 1 8 pp.
Barona de la O, F., 1 951 , “Alkali-Aggregate Expansion Corrected with Portland-Slag Cement,” ACI Journal
Proceedings , V. 47, No. 3, Mar., pp. 545-552.
Bassuoni, M. T., and Nehdi, M. L., 2009, “Durability of
Self-Consolidating Concrete to Different Exposure Regimes
of Sodium Sulfate Attack,” Materials and Structures , V. 42,
No. 8, Oct., pp. 1 039-1 057. doi: 1 0.1 61 7/s1 1 527-008-9442-2
Bastiensen, R.; Mourn, J.; and Rosenquist, I., 1 957, “Some
Investigations o f Alum Shale in Construction (Bidragfl
Belysning av visse Bygningstekniske Problemer ved
Osloomradets Alunskifere),” Publication No.22 , Norwegian
Geotechnical Institute, Oslo, Norway, 69 pp. (in Norwegian)
Bates, P. H., and Klein, A. A., 1 91 7, “Properties of the
Calcium Silicates and Calcium Aluminate Occurring in
Normal Portland Cement,” Technologic Papers o f the
Bureau o fStandards , U.S. Department of Commerce, Washington, DC., No. 78, June, 52 pp.
Bates, P. H.; Phillips, A. J.; and Wig, R. J., 1 91 3, “Action
of the Salts in Alkali Water and Sea Water on Cement,” Technologic Papers o fthe Bureau o fStandards , U.S. Department
of Commerce, Washington, DC., No. 1 2, GPO, 1 57 pp.
Bavarian, B., and Reiner, L., 2004, “Improving Durability
of Reinforced Concrete Structures using Migrating Corrosion Inhibitors,” Paper 04323, Proceedings, Corrosion 2004 ,
National Association of Corrosion Engineers (NACE), New
Orleans, LA, pp. 04323/1 -04323/11 .
Baweja, D.; Sirivivatnanon, V.; Gross, W.; and Laurie, G.,
1 994, “High-Performance Australian Concretes for Marine
Applications,” High-Per formance Concrete , Proceedings
of the Second ACI International Conference, SP-1 49, V.
M. Malhotra, ed., American Concrete Institute, Farmington
Hills, MI, pp. 363-378.
Beaudoin, J. J., and Brown, P. W., 1 992, “The Structure
of Hardened Cement Paste,” Proceedings o f the 9th International Congress on the Chemistry o f Cement, V. I, New
Delhi, National Council for Cement and Building Materials,
pp. 485-525.
Bellmann, F., and Stark, J., 2007, “Prevention of Thaumasite Formation in Concrete Exposed to Sulfate Attack,”
Cement and Concrete Research , V. 37, No. 8, Aug., pp.
1 21 5-1 222. doi: 1 0.1 01 6/j.cemconres.2007.04.007
Bellmann, F., and Stark, J., 2008, “The Role of Calcium
Hydroxide in the Formation of Thaumasite,” Cement and
Concrete Research , V. 38, No. 1 0, Oct., pp. 1 1 54-1 1 61 . doi:
1 0.1 01 6/j.cemconres.2008.04.005
Bellport, B. P., 1 968, “Combating Sulphate Attack on
Concrete on Bureau of Reclamation Projects,” Performance
o f Concrete , University of Toronto Press, pp. 77-92.
Bentz, D. P., 2000, “Fibers, Percolation, and Spalling of
High Performance Concrete,” ACI Materials Journal , V. 97,
No. 3, May-June, pp. 351 -359.
Bentz, D. P.; Garboczi, E. J.; and Snyder, K. A., 1 999, “A
Hard Core/Soft Shell Microstructural Model for Studying
Percolation and Transport in Three-Dimensional Composite
63
Media,” NISTIR 6265 , U.S. Department of Commerce,
Washington, DC, Jan.
Bentz, D. P.; Schlangen, E.; and Garboczi, E. J., 1 995,
“Computer Simulation of Interfacial Zone Microstructure
and Its Effect on the Properties of Cement-Based Composites,” Materials Science o f Concrete IV, J. Skalny and S.
Mindess, eds., American Ceramic Society, Westerville, OH,
pp. 1 55-1 99.
Berke, N. S., 1 985, “Effects of Calcium Nitrite and Mix
Design on the Corrosion Resistance of Steel in Concrete
(Part l),” NACE Corrosion 85 , Paper No.273, National Association of Corrosion Engineers, Houston, TX.
Berke, N. S., 1 987, “Effect of Calcium Nitrite and Mix
Design on the Corrosion Resistance of Steel in Concrete
(Part 2, Long-Term),” NACE Corrosion 87, Paper No. 1 32,
National Association of Corrosion Engineers, Houston, TX.
Berke, N. S., 1 988, “Microsilica and Concrete Durability,” Transportation Research Record 1204 , Transportation Research Board, Washington, DC, pp. 21 -26.
Berke, N. S.; Dallaire, M. P.; and Hicks, M. C., 1 992,
“Plastic, Mechanical, Corrosion, and Chemical Resistance
Properties of Silica Fume (Microsilica) Concretes,” Fly
Ash, Silica Fume, Slag, and Natural Pozzolans in Concrete:
Fourth CANMET/ACI International Con ference , V. M.
Malhotra, ed., American Concrete Institute, Farmington
Hills, MI, pp. 11 25-1 1 50.
Berke, N. S., and Hicks, M. C., 1 992, “Estimating the
Life Cycle of Reinforced Concrete Decks and Marine Piles
Using Laboratory Diffusion and Corrosion Data,” Corrosion
Forms and Control for In frastructure , V. 1 1 39, pp. 207-231 .,
207-225. doi: 1 0.1 520/STP1 9764S
Berke, N. S.; Hicks, M. C.; and Hoopes, R. J., 1 994,
“Condition Assessment of Field Structures with Calcium
Nitrite,” Concrete Bridges in Aggressive Environments:
Philip D. Cady International Symposium , SP-1 51 , R. E.
Weyers, ed., American Concrete Institute, Farmington Hills,
MI, pp. 43-72.
Berke, N. S., and Roberts, L. R., 1 989, “Use of Concrete
Admixtures to Provide Long-Term Durability from Steel
Corrosion,” Superplasticizers and Other Chemical Admix tures in Concrete , Proceedings of the Third CANMET/
ACI International Conference, SP-11 9, V. M. Malhotra,
ed., American Concrete Institute, Farmington Hills, MI, pp.
383-403.
Berke, N. S., and Rosenberg, A., 1 989, “Technical Review
of Calcium Nitrite Corrosion Inhibitor in Concrete,” Trans portation Research Record 1211 , Transportation Research
Board, Washington, DC, 1 8 pp.
Berke, N. S.; Scali, M. J.; Regan, J. C.; and Shen, D. F.,
1 991 , “Long-Term Corrosion Resistance of Steel in Silica
Fume and/or Fly Ash Containing Concretes,” Durability o f
Concrete , Proceedings of the Second CANMET/ACI International Conference, SP-1 26, V. M. Malhotra, ed., American
Concrete Institute, Farmington Hills, MI, pp. 393-422.
Bérubé, M.-A., and Duchesne, J., 1 992, “Evaluation
of Testing Methods Used for Assessing the Effectiveness
of Mineral Admixtures in Suppressing Expansion Due to
Alkali-Aggregate Reaction,” Fly Ash, Silica Fume, Slag,
64
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
and Natural Pozzolans in Concrete ,
Proceedings of the
Fourth CANMET/ACI International Conference, SP-1 32, V.
M. Malhotra, ed., American Concrete Institute, Farmington
Hills, MI, pp. 549-575.
Bérubé, M.-A.; Duchesne, J.; Dorion, J.; and Rivest, M.,
2002, “Laboratory Assessment of Alkali Contribution by
Aggregates to Concrete and Application to Concrete Structures Affected by Alkali-Silica Reactivity,” Cement and
Concrete Research , V. 32, No. 8, Aug., pp. 1 21 5-1 227. doi:
1 0.1 01 6/S0008-8846(02)00766-4
Bérubé, M.-A., and Fournier, B., 1 992a, “Accelerated
Test Methods for Alkali-Aggregate Reactivity,” Advances in
Concrete Technology , V. M. Malhotra, ed., CANMET/EMR,
Ottawa, ON, Canada, pp. 583-627.
Bérubé, M.-A., and Fournier, B., 1 992b, “Effectiveness
of the Accelerated Mortar Bar Method, ASTM C9 Proposal
P21 4 or NBRI, for Assessing Potential AAR in Quebec
(Canada),” Proceedings o f the Ninth International Con ference on Alkali-Aggregate Reaction in Concrete , Concrete
Society, Slough, UK, pp. 92-1 01 .
Bérubé, M.-A., and Fournier, B., 1 993, “Testing for
Alkali-Aggregate Reactivity in Concrete,” Memoria del
Seminario Internacional sobre Tecnología del Concreto ,
R. Rivera Villareal, ed., Universidad Autónoma de Nuevo
León, Nuevo León, Mexico, Oct., pp. 54-78.
Bessey, G. E., and Lea, F. M., 1 953, “The Distribution of
Sulphates in Clay Soils and Ground Waters,” Proceedings Institution o f Civil Engineers , V. 2, No. 2, pp. 1 59-1 81 . doi:
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1 0.1 680/iicep.1 953.11 030
Bhatty, M. S. Y., and Greening, N. R., 1 978, “Interaction
of Alkalis with Hydrating and Hydrated Calcium Silicates,”
Proceedings o f the Fourth International Con ference on the
Effects o f Alkalis in Cement and Concrete , Purdue Univer-
sity, West Lafayette, IND, pp. 87-11 2.
Bickley, J. A.; Hemmings, R. T.; Hooton, R. D.; and
Balinsky, J., 1 994, “Thaumasite Related Deterioration of
Concrete Structures,” Proceedings o f Concrete Technology:
Past, Present and Future , SP-1 44, American Concrete Institute, Farmington Hills, MI, Mar., pp. 1 59-1 75.
Biczok, I., 1 967, Concrete Corrosion and Concrete
Protection , Chemical Publishing Co., Inc, New York.
Biczok, I., 1 972, Concrete Corrosion–Concrete Protec tion , eighth edition, Akademiai Kiado, Budapest, 545 pp.
Bier, T. A., 1 987, “Inf uence o f Type o f Cement and
Curing on Carbonation Progress and Pore Structure of
Hydrated Cement Pastes,” Microstructural Development
During Hydration o f Cement, Proceedings of the Materials
Research Society, Pittsburgh, PA, V. 85, pp. 1 23-1 34.
Bier, T. A.; Ludirdja, D.; Young, J.; and Berger, R.,
1 988, “The Effect of Pore Structure on the Permeability of
Concrete,” Pore Structure and Permeability o f Cementitious
Materials , Proceedings of the Materials Research Society,
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Bijen, J., and van Selst, R., 1 991 , “Effects of Fly Ash on
Carbonation of Concrete with Portland Blast-Furnace Slag
Cement,” Durability o f Concrete , Proceedings of the Second
CANMET/ACI International Conference, SP-1 26, V. M.
Malhotra, ed., American Concrete Institute, Farmington
Hills, MI, pp. 1 001 -1 030.
Blanks, R. F., 1 950, “Fly Ash as a Pozzolan,” ACI Journal
Proceedings , V. 47, No. 9, Sept., pp. 701 -707.
Bleszynski, R. F., 2002, “The Performance and Durability
of Concrete with ternary Blends of Silica Fume and BlastFurnace Slag,” PhD thesis, University of Toronto.
Bleszynski, R. F.; Hooton, R. D.; Thomas, M. D. A.; and
Rogers, C. A., 2002, “Durability of Ternary Blend Concretes
with Silica Fume and Blastfurnace Slag: Laboratory and
Outdoor Exposure Site Studies,” ACI Materials Journal , V.
99, No. 5, Sept.-Oct., pp. 499-508.
Bleszynski, R. F., and Thomas, M. D. A., 1 998, “Microstructural Studies of Alkali-Silica Reaction in Fly Ash
Concrete Immersed in Alkaline Solutions,” Advanced
Cement Based Materials , V. 7, No. 2, Mar., pp. 66-78. doi:
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Bobrowski, G., and Youn, D. J., 1 993, “Corrosion Inhibitors in Cracked Concrete: An Admixture Solution,” Concrete
2000: Economic and Durable Construction through Excellence , Proceedings of the International Conference, R. K.
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Boddy, A. M.; Hooton, R. D.; and Thomas, M. D. A.,
2003, “The Effect of the Silica Content of Silica Fume on
its Ability to Control Alkali-Silica Reaction,” Cement and
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Standard
Sharing
our chatsJ.; and Baumann,
Bognacki,
C. J.;Group
Pirozzi, and
M.; Marsanno,
W. C., 201 0, “Rapid Chloride Permeability Testing’s Suitability for Use in Performance-Based Specifcations,”
Concrete International , V. 28, No. 5, May, pp. 47-52.
Bogue, R. H., 1 955, The Chemistry o f Portland Cement ,
second edition, Reinhold Publishing Corporation, New York.
Branca, C.; Fratesi, R.; Moriconi, G.; and Simoncini,
S., 1 992, “In f uence o f Fly Ash on Concrete Carbonation
and Rebar Corrosion,” Fly Ash, Silica Fume, Slag, and
Natural Pozzolans in Concrete , Proceedings of the Fourth
CANMET/ACI International Conference, SP-1 3 2, V. M.
Malhotra, ed., American Concrete Institute, Farmington
Hills, MI, pp. 245-256.
Broomfeld, J., 2007, Corrosion o f Steel in Concrete:
Understanding, Investigation and Repair, second edition,
Taylor & Francis, London, 296 pp.
Brown, P. V., and Bothe Jr., J. V., 2004, “The System
CaO-Al 2 O 3 -CaCl 2 -H 2 O at 23±2°C and Mechanisms of
Chloride Binding in Concrete,” Cement and Concrete
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Brown, P. W., 2002, “Thaumasite Formation and Other
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S0958-9465(01 )00081 -6
Brown, P. W.; Taylor, H. F. W.; Young, J. F.; and Johannsen,
V., 1 986, “The Hydration of Tricalcium Aluminate and Tetracalcium Aluminoferrite in the Presence of Calcium Sulfate,”
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
Materials and Structures , V.
1 9, No. 2, Mar., pp. 1 37-1 47.
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Browne, R. D., 1 980, “Mechanisms of Corrosion of Steel
in Concrete in Relation to Design, Inspection, and Repair of
Offshore and Coastal Structures,” Performance o f Concrete
in Marine Environment , SP-65, V. M. Malhotra, ed., American Concrete Institute, Farmington Hills, MI, pp. 1 69-204.
Brunette, B. E., and Lundy, J. R., 1 996, “Use and Effects
of Studded Tires on Oregon Pavements,” Transportation
Research Record, V. 1 536, pp. 64-72. doi: 1 0.31 41 /1 536-1 0
Buck, A. D.; Houston, B. J.; and Pepper, L., 1 953, “Effectiveness of Mineral Admixtures in Preventing Excessive
Expansion of Concrete Due to Alkali-Aggregate Reaction,”
ACI Journal Proceedings , V. 50, No. 1 0, Oct., p. 1 1 60.
Buenfeld, N. R., and Newman, J. B., 1 986, “The Development and Stability of Surface Layers on Concrete Exposed
to Sea-Water,” Cement and Concrete Research , V. 1 6, No.
5, Sept., pp. 721 -732. doi: 1 0.1 01 6/0008-8846(86)90046-3
Buenfeld, N. R., and Newman, J. B., 1 987, “Examination
of Three Methods for Studying Ion Diffusion in Cement
Pastes, Mortars and Concrete,” Materials and Structures , V.
20, No. 1 , Jan., pp. 3-1 0. doi: 1 0.1 007/BF02472720
Buenfeld, N. R.; Glass, G. K.; Hassanein, A. M.; and
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Highway Administration, Washington, DC.
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Coussy, O., and Monteiro, P. J. M., 2008, “Poroelastic
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GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
Duchesne, J., and Bérubé, M. A., 1 992, “An Autoclave
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Reinforcing Steel in Concrete Exposed to a Marine Environment,” Technical Note No. N-1 032, U.S. Naval Civil Engineering Laboratory, Port Hueneme, CA, July, 42 pp.
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Blast Furnace Slag Concrete,” Nordic Concrete Research
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(Norsk Betongforening ), Nordic Concrete Federation, Oslo,
Norway, No. 1 1 , Feb., pp. 55-67.
Hall, C., and Hoff, W. D., 201 2, Water Transport in Brick,
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370 pp.
Hallet, B.; Walder, J. S.; and Stubbs, C. W., 1 991 , “Weathering by Segregation Ice Growth in Microcracks at Sustained
S ub- Zero Temperatures: Verifcation from an Experimental
Study Using Acoustic Emissions,” Perma frost and Perigla cial Proceedings , V. 2, 283-300 pp.
Hamilton, J. J., and Handegord, G. O., 1 968, “The Performance of Ordinary Portland Cement Concrete in Prairie
Soils of High Sulphate Content,” Performance o fConcrete—
Resistance o fConcrete to Sulphate and Other Environmental
Conditions , Thorvaldson Symposium, University of Toronto
Press, Toronto, ON, Canada, pp. 1 35-1 58.
Hansen, W. C., 1 944, “Studies Relating to the Mechanism
by Which the Alkali-Silica Reaction Proceeds in Concrete,”
ACI Journal Proceedings , V. 41 , pp. 21 3-227.
Hansen, W. C., 1 960, “Inhibiting Alkali-Aggregate Reaction with Barium Salts,” ACI Journal Proceedings , V. 57,
No. 9, Mar., pp. 881 -883.
Hansen, W. C., and Pressler, E. E., 1 947, “Solubility of
Ca(OH) 2 and CaSO 4 .2H 2 O in Dilute Alkali Solutions,”
Industrial & Engineering Chemistry , V. 39, No. 1 0, Oct., pp.
1 280-1 282. doi: 1 0.1 021 /ie50454a005
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Materials Journal, V. 89, No. 3, May-June, pp. 238-241 .
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Oct., pp. 1 -20.
Hayes, C. F., 2007, “Test Method for Water-Soluble
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Haynes, H.; O’ Neill, R.; and Mehta, P. K., 1 996, “Concrete
Deterioration from Physical Attack by Salts,” Concrete
International , V. 1 8, No. 1 , Jan., pp. 63-68.
Haynes, H.; O’Neill, R.; Neff, M.; and Mehta, P. K., 2008,
“Salt Weathering Distress on Concrete Exposed to Sodium
Sulfate Environment,” ACI Materials Journal, V. 1 05, No. 1 ,
Jan.-Feb., pp. 35-43.
Haynes, H.; O’Neill, R.; Neff, M.; and Mehta, P. K., 201 0,
“Salt Weathering of Concrete by Sodium Carbonate and
Sodium Chloride,” ACI Materials Journal , V. 1 07, No. 3,
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Hearn, N.; Hooton, R. D.; and Nokken, M. R., 2006, “Pore
Structure and Permeability,” Signifcance o f Tests and Prop erties o f Concrete and Concrete Making Materials ,” ASTM
STP 1 69D, pp. 238-252.
Heinz, D., and Ludwig, U., 1 987, “Mechanism of
Secondary Ettringite Formation in Mortars and Concretes
Subjected to Heat Treatment,” Concrete Durability: Kath erine and Bryant Mather International Con ference , J. M.
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Helmuth, R. A., 1 960a, “Capillary Size Restrictions on
Ice Formation in Hardened Portland Cement Pastes,” Fourth
International Symposium on the Chemistry o fCement, NBS,
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Helmuth, R. A., 1 960b, “Capillary Size Restrictions on Ice
Formation in Hardened Portland Cement Pastes,” Proceed-
ings o f the Fourth International Symposium on the Chem istry on Cement, Monograph No. 43, National Bureau of
Standards, Washington, DC, V. 2, pp. 855-869.
Helmuth, R. A., 1 960c, “Frost Action in Concrete”
Proceedings o f the Fourth International Symposium on the
Chemistry o f Cement, Monograph No. 43, National Bureau
of Standards, Washington, DC, V. 2, pp. 829-833.
Henriksen, H.; Kjaer, U.; and Lundberg, L., 2000,
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for Tunnel Elements and Ramps and Experiences from the
Concrete Production and Castings,” Øresund Link Immersed
Tunnel Conference, Copenhagen, Apr., pp. D1 -1 - D1 -1 2.
Hewlett, P. C., 1 998, Lea’s Chemistry o f Cement and
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Hill, J.; Byars, E. A.; Sharp, J. H.; Lynsdale, C. J.; Cripps,
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Hobbs, D. W., 1 989, “Effect of Mineral and Chemical
Admixtures on Alkali-Aggregate Reaction,” Proceedings o f
the 8th International Con ference on Alkali-Aggregate Reac tion , K. Okada, S. Nishibayashi, and M. Kawamura, eds.,
E&FN Spon, London, UK, pp. 1 73-1 86.
Holland, T. C., and Gutschow, R. A., 1 987, “Erosion
Resistance with Silica-Fume Concrete,” Concrete Interna tional, V. 9, No. 3, Mar., pp. 32-40.
Holm, J., 1 987, “Comparison of the Corrosion Potential of
Calcium Chloride and a Calcium Nitrite Based Non-Chloride
Accelerator—A Macro-Cell Corrosion Approach,” Corro sion, Concrete, and Chlorides , SP-1 02, F. W. Gibson, ed.,
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Test for Evaluation of Concrete Quality,” Pore Structure and
Permeability o f Cementitious Materials , Materials Research
Society Symposium Proceedings, V. 1 37, L. R. Roberts and
J. P. Skalny, eds., pp. 1 41 -1 49.
Hooton, R. D., 1 991 , “New Aggregates Alkali-Reactivity
Test Methods,” Research Report MAT-91-14 , Ministry of
Transportation, ON, Canada.
Hooton, R. D . , 1 9 9 3 , “In f uence o f S ilica F ume Replace ment of Cement on Physical Properties and Resistance to
Sulfate Attack, Freezing and Thawing, and Alkali-Silica
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Recent Developments,” Proceedings o f the RILEM Work-
shop on Performance o fCement-Based Materials in Aggres sive Aqueous Environments , Ghent, Belgium, pp. 75-77.
Hooton, R. D., 2008, “Bridging the Gap Between Research
and Standards,” Cement and Concrete Research , V. 38, No.
2, Feb., pp. 247-258. doi: 1 0.1 01 6/j.cemconres.2007.09.01 2
Hooton, R. D., and Brown, P. W., 2009, “Development
of Test Methods to Address the Various Mechanisms of
Sulfate Attack,” RILEM Proceedings on PRO63, Concrete
in Aggressive, Aqueous Environments , Toulouse, V. 2, pp.
280-297.
Hooton, R. D., and Rogers, C. A., 1 989, “Evaluation of
Rapid Test Methods for Detecting Alkali-Reactive Aggregates,” Proceedings o f the 8th International Con ference
on Alkali-Aggregate Reaction , The Society of Materials
Science, Japan, pp. 439-444.
Hooton, R. D., and Rogers, C. A., 1 992, “Development of
the NBRI Rapid Mortar Bar Test Leading to Its Use in North
America,” Proceedings o f the 9th International Con ference
on Alkali-Aggregate Reaction in Concrete , The Concrete
Society, Weham, Slough, V. 1 , pp. 461 -467.
Hooton, R. D., and Thomas, M. D. A., 2002, “The Use of
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of Sulfate Attack,” PCA R&D Serial No. 2658 , Portland
Get more
standards from
Cement Association, Skokie,
IL, 1 0FREE
pp.
Hooton, R. D.; Geiker, M. R.; and Bentz, E. C., 2002,
“Effects of Curing on Chloride Ingress and Implications for
Service Life,” ACI Materials Journal , V. 99, No. 2, Mar.Apr., pp. 201 -206.
Hooton, R. D.; Ramezanianpour, A.; and Schutz, U.,
201 0, “Decreasing the Clinker Component in Cementing
Materials: Performance of Portland-Limestone Cements in
Concrete in combination with Supplementary Cementing
Materials,” CD Proceedings, 2010 Concrete Sustainability
Con ference , National Ready Mixed Concrete Association,
Silver Springs, MD, 1 5 pp.
Hong, S.-Y., and Glasser, F. P., 1 999, “Alkali Binding
in Cement Pastes: Part I. The C-S-H Phase,” Cement and
Concrete Research , V. 29, No. 1 2, Dec., pp. 1 893-1 903. doi:
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Hong, S.-Y., and Glasser, F. P., 2002, “Alkali Sorption
by C-S-H and C-A-S-H Gels: Part II. Role of Alumina,”
Cement and Concrete Research , V. 32, No. 7, July, pp. 11 01 11 11 . doi: 1 0.1 01 6/S0008-8846(02)00753-6
Hope, B. B.; Ip, A. K.; and Manning, D. G., 1 985, “Corrosion and Electrical Impedance in Concrete,” Cement and
Concrete Research , V. 1 5, No. 3, May, pp. 525-534. doi:
1 0.1 01 6/0008-8846(85)901 27-9
Horiguchi, K.; Chosokabi, T.; Ikabata, T.; and Suzuki, Y.,
1 994, “Rate of Carbonation in Concrete Made with Blended
Cement,” Durability o f Concrete , Proceedings of the Third
CANMET/ACI International Conference, SP-1 45, V. M.
Malhotra, ed., American Concrete Institute, Farmington
Hills, MI, pp. 91 7-932.
Hudec, P. P., and Larbi, E. Y., 1 989, “Chemical Treatments
and Additives to Minimize Alkali Reactivity,” Proceedings
o f the 8th International Con ference on Alkali-Aggregate
Reaction , The Society of Materials Science, Japan, pp.
1 93-1 98.
Hurst, W. D., 1 968, “Experience in the Winnipeg Area
with Sulphate-Resisting Cement Concrete,” Performance o f
Concrete-Resistance o fConcrete to Sulphate and Other Envi ronmental Conditions , Thorvaldson Symposium, University
of Toronto Press, Toronto, ON, Canada, pp. 1 25-1 34.
Hussain, S. E., and Rasheeduzzafar, 1 994, “Corrosion Resistance Performance of Fly Ash Blended Cement
Concrete,” ACI Materials Journal , V. 91 , No. 3, May-June,
pp. 264-272.
Ibrahim, A. A.; Abuazza, O. A.; and Tarrani, F. A., 1 997,
“Prestressed Concrete Cylinder Pipes Exposed Internally to
Aggressive Water,” Durability o f Concrete , SP-1 70, V. M.
Malhotra, ed., American Concrete Institute, Farmington
Hills, MI, pp. 437-439.
Ideker, J. H.; East, B. L.; Folliard, K. J.; Thomas, M.
D. A.; and Fournier, B., 201 0, “The Current State of the
Accelerated Concrete Prism Test,” Cement and Concrete
Research , V. 40, No. 4, Apr., pp. 550-555. doi: 1 0.1 01 6/j.
cemconres.2009.08.030
Ideker, J. H.; Folliard, K. J.; Juenger, M. C. G.; and
Bentivegna, A. F., 201 2, “Do Current Laboratory Test
Methods Accurately Predict Alkali-Silica Reactivity?” ACI
Materials Journal, V. 1 09, No. 4, July-Aug., pp. 395-400.
Standard
Group
and
chats
Idorn,Sharing
G. M., and
Roy, D.
M.,our
1 986,
“Opportunities with
Alkalies in Concrete Testing, Research, and Engineering
Practice,” Alkalies in Concrete , STP-930, ASTM International, West Conshohocken, PA, Jan., pp. 5-1 5.
Irassar, E. F.; Di Maio, A.; and Batic, O. R., 1 996, “Sulfate
Attack on Concrete with Mineral Admixtures,” Cement and
Concrete Research , V. 26, No. 1 , Jan., pp. 11 3-1 23. doi:
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Janssen, D. J., 2001 , “Highway Concrete Pavement Technology Volume II: Field Evaluation of SHRP C-203 (FreezeThaw Resistance) Test Sites,” Federal Highway Administration, Washington, DC.
Janssen, D. J., and Snyder, M. B., 1 993, “Mass Loss Experience with ASTM C666: With and Without Deicing Salt,”
Proceedings, International Workshop on the Resistance
o f Concrete to Scaling Due to Freezing in the Presence o f
Deicing Salts , Laval University, Sainte-Foy, QC, Canada,
E&FN Spon, London, Aug., pp. 1 37-1 51 .
Janssen, D. J., and Snyder, M. B., 1 994, “Resistance of
Concrete to Freezing and Thawing,” SHRP-C-391, Strategic
Highway Research Program, Federal Highway Administration, Washington, DC.
Janssen, D. J.; DuBose, J. D.; Patel, A. J.; and Dempsey,
B. J., 1 986, “Predicting the Progression of D-Cracking,”
Transportation Engineering Series No. 44, University of
Illinois, Champaign, IL.
Janssen, D. J.; Dyer, R. M.; and Elkey, W. E., 1 995, “Effect
of Pumping on Entrained Air Voids: Role of Pressure,”
Proceedings, CONSEC 95, Sakai, K., N. Banthia, and O. E.
Gjørv, eds., Sapporo, Japan, E&FN Spon, Tokyo, pp. 233-242.
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
Johannesson, B. F., 2003, “A Theoretical Model
Describing Diffusion of a Mixture of Different Types of Ions
in Pore Solution of Concrete Coupled to Moisture Transport,” Cement and Concrete Research , V. 33, No. 4, Apr.,
pp. 481 -488. doi: 1 0.1 01 6/S0008-8846(02)00993-6
Johansen, V.; Thaulow, N.; and Skalny, J., 1 993, “Simultaneous Presence of Alkali-Silica Gel and Ettringite in
Concrete,” Advances in Cement Research , V. 5, No. 1 7, Jan.,
pp. 23-29. doi: 1 0.1 680/adcr.1 993.5.1 7.23
Julio-Betancourt, G. A., and Hooton, R. D., 2005, “Effect
of De-Icer and Anti-Icer Chemicals on the Durability,
Microstructure and Properties of Cement-Based Materials,”
Proceedings o f the Canadian Society o f Civil Engineering ,
Toronto, ON, Canada, pp. 2669.
Jus tnes, H. , and Nygaard, E. C . , 1 9 9 4, “The In f uence o f
Technical Calcium Nitrate Additions on the Chloride Binding
Capacity of Cement and the Rate of Chloride Induced
Corrosion of Steel Embedded in Mortars,” Proceedings o f
the International Con ference on Corrosion and Corrosion
Protection o f Steel in Concrete
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Kalousek, G. L.; Porter, L. C.; and Benton, E. J., 1 972,
“Concrete for Long-Time Service in Sulfate Environment,”
Cement and Concrete Research , V. 2, No. 1 , Jan., pp. 79-89.
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Kalousek, G. L.; Porter, L. C.; and Harboe, E. J., 1 976,
“Past, Present, and Potential Developments of SulphateResisting Concretes,” Journal o f Testing and Evaluation , V.
4, No. 5, Sept., pp. 347-354. doi: 1 0.1 520/JTE1 0522J
Kaneuji, M.; Winslow, D. N.; and Dolch, W. L., 1 980,
“The Relationship Between an Aggregate’s Pore Size Distribution and Its Freeze Thaw Durability in Concrete,” Cement
and Concrete Research , V. 1 0, No. 3, May, pp. 433-441 . doi:
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Kauer, J. A., and Freeman, R. L., 1 955, “Effect of Carbon
Dioxide on Fresh Concrete,” ACI Journal Proceedings , V.
52, No. 4, Dec., pp. 447-454.
Kawamura, M., and Takemoto, K., 1 988, “Correlation
between Pore Solution Composition and Alkali Silica Expansion in Mortars Containing Various Fly Ashes and Blastfurnace Slags,” International Journal o f Cement Composites
and Lightweight Concrete , V. 1 0, No. 4, pp. 21 5-223. doi:
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Kawamura, M.; Takemoto, K.; and Hasaba, S., 1 987,
“Effectiveness of Various Silica Fumes in Preventing
Alkali-Silica Expansion,” Concrete Durability: Katharine
and Bryant Mather International Con ference , SP-1 00, J.
M Scanlon, ed., American Concrete Institute, Farmington
Hills, MI, pp. 1 809-1 820.
Kelham, S., 1 996, “The Effect of Cement Composition
and Fineness on Expansion Associated with Delayed Ettringite Formation,” Cement and Concrete Composites , V. 1 8,
No. 3, pp. 1 71 -1 79. doi: 1 0.1 01 6/0958-9465(95)0001 3-5
Kennerley, R. A., 1 988, “Experience in New Zealand with
Pozzolans, Fly Ash and Slag,” Concrete 88 Workshop , W.
G. Ryan, ed., Concrete Institute of Australia, Sydney, pp.
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Kennerley, R. A.; St. John, D. A.; and Smith, L. M., 1 981 ,
“A Review of Thirty Years of Investigation of the AlkaliAggregate Reaction in New Zealand,” Proceedings o f the
5th International Con ference on Alkali-Aggregate Reaction ,
Cape Town, CSIRO, Pretoria, Paper S252/1 2.
Kernes, R. G.; Edwards, J. R.; Dersch, M. S.; Lange, D.
A.; and Barkan, C. P. L., 201 1 , “Investigation of the Impact
of Abrasion as a Concrete Crosstie Rail Seat Deterioration
(RSD) Mechanism,” AREMA 2011 Annual Con ference in
Conjunction with Railway Interchange 2011 , American
Railway Engineering and Maintenance-of-Way Association,
Lanham, MD, 24 pp.
Keshari, S., 2009, “Effect of Constituent Materials and
Curing Methods on the Abrasion Resistance and Durability
of High Performance Concrete for Pre-Cast Pre-Stressed
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Kevern, J. T.; Schaefer, V. R.; and Wang, K., 2009, “The
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Resistance,” Journal o f Testing and Evaluation , V. 37, No.
4, pp. 1 -6.
Keyser, J. H., 1 971 , “Resistance of Various Types of Bituminous Concrete and Cement Concrete to Wear by Studded
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Kjellsen, K. O.; Detwiler, R. J.; and Gjørv, O. E.,
1 991 , “Development of Microstructures in Plain Cement
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Concrete Research , V. 21 , No. 1 , Jan., pp. 1 79-1 89. doi:
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Klieger, P., 1 952, “Studies of the Effect of Entrained Air
on the Strength and Durability of Concretes Made with
Various Sizes of Aggregate,” Research Department Bulletin
RX040 , Portland Cement Association, Skokie, IL.
Klieger, P., 1 956, “Further Studies on the Effect of
Entrained Air on Strength and Durability of Concrete with
Various Sizes of Aggregates,” Research Department Bulletin
RX077, Portland Cement Association, Skokie, IL.
Klieger, P., and Gebler, S., 1 987, “Fly Ash and Concrete
Durability,” Concrete Durability: Katharine and Bryant
Mather International Con ference , SP-1 00, J. M. Scanlon,
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1 043-1 069.
Klieger, P., and Greening, N. R., 1 969, “Properties of
Expansive Cement Concretes,” Proceedings o f the 5th International Symposium on the Chemistry o f Cement, Tokyo,
Japan, pp. 439-456.
Klieger, P.; Monfore, G.; Stark, D.; and Teske, W., 1 974,
“D-Cracking of Concrete Pavements in Ohio,” Final Report ,
Ohio-DOT-1 1 -74.
Kollek, J. J.; Varma, S. P.; and Zaris, C., 1 986, “Measurement of OH – Concentrations of Pore Fluids and Expansion
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Kompen, R., 1 994, “Prosjektering for bestandighet
(Development of Resistance),” Prosjektering og Produksjon
av bestandige betongkonstruksjoner (Design and Produc tion o f Resistant Concrete Structures) , Jan. 1 0-1 2, 1 994,
Trondheim, Norsk Betonforening. (in Norwegian)
Krauskopf, K. B., and Bird, D. K., 1 995, Introduction to
Geochemistry , third edition, Stanford University Press, CA,
647 pp.
Krukar, M., and Cook, J. C., 1 973, “Effect of Studded Tires
on Various Pavements and Surfaces,” Highway Research
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Kummer, H. W., and Meyer, W. E., 1 967, “Tentative
Skid-Resistance Requirements for Main Rural Highways,”
NCHRP Report No. 37, Highway Transportation Research
Board, Washington, DC, 80 pp.
Lane, D. S., 1 987, “Long-Term Mortar Bar Expansion
Tests for Potential Alkali-Aggregate Reactivity,” Proceed-
ings o f the 7th International Con ference on Concrete AlkaliAggregate Reactions , P. E. Grattan-Bellew, ed., Noyes, Park
Ridge, NJ, pp. 336-341 .
Lawrence, C. D., 1 995, “Delayed Ettringite Formation:
An Issue?” Materials Science o f Concrete IV, J. Skalny and
S. Mindess, eds., American Ceramic Society, Westerville,
OH, pp. 1 1 3-1 54.
Lawrence, M., and Vivian, H. F., 1 961 , “The Reactions
Get
moreAustralian
FREE standards
from
of Various Alkalies with
Silica,”
Journal Applied
Science , Commonwealth Scientifc and Industrial Research
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Lea, F. M., 1 971 , The Chemistry o f Cement and Concrete ,
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Lerch, W.; Ashton, F. W.; and Bogue, R. H., 1 929, “The
Sulphoaluminates of Calcium,” Bureau o f Standards
Journal o f Research , V. 2, No. 4, pp. 71 5-731 . doi: 1 0.6028/
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Lichtner, P. C.; Steefel, C. I.; and Oelkers, E. H., 1 996,
“Reactive Transport in Porous Media,” Reviews in Mineralogy , P. H. Ribbe, ed., Mineralogical Society of America,
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Litvan, G. G., 1 972, “Phase Transitions of Adsorbates:
IV, Mechanism of Frost Action in Hardened Cement Paste,”
Journal o f the American Ceramic Society , V. 55, No. 1 , Jan.,
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Litvan, G. G., 1 978, “Adsorption Systems atTemperatures
Below the Freezing Point of the Adsorptive,” Advanced
Colloidal and Interfacial Science , V. 9, No. 4, June, pp.
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Liu, T., 1 994, “Abrasion Resistance,” Signifcance o f Tests
and Properties o fConcrete and Concrete-Making Materials ,
STP-1 69C, P. Klieger and J. Lamond, eds., ASTM International, West Conshohocken, PA, pp. 1 82-1 92.
Liu, Y.; Yen, T.; and Hsu, T., 2006, “Abrasion Erosion
of Concrete by Water-Borne Sand,” Cement and Concrete
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Longuet, P., 1 976, “La protection des armatures dans le
beton armé elaboréavec des ciments de laitier,” Silicates
Industrials , V. 7, No. 8, pp. 321 -328.
Lossing, F. A., 1 966, “Sulfate Attack on Concrete Pavements in Mississippi,” Highway Research Record No. 113 ,
Highway Transportation Research Board, Washington, DC,
pp. 88-1 02.
Lu, D.; Fournier, B.; Grattan-Bellew, P.; Xu, Z.; and Tang,
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Terzaghi, R. D., 1 949, “Concrete Deterioration Due to
Carbonic Acid,” Journal o f the Boston Society o f Civil Engi neers , V. 36, No. 2, Apr., pp. 1 36-1 60.
Thaumasite Expert Group, 1 999, “The Thaumasite Form
of Sulfate Attack: Risks, Diagnosis, Remedial Works and
Guidance on New Construction,” Report o f the Thaumasite
Expert Working Group , DETR, London, UK, 1 80 pp.
Thomas, M. D. A., 1 996, “Field Studies of Fly Ash
Concrete Structures Containing Reactive Aggregates,”
Magazine o f Concrete Research , V. 48, No. 1 77, Dec., pp.
265-279. doi: 1 0.1 680/macr.1 996.48.1 77.265
80
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
Thomas, M. D. A., 1 997, “Laboratory and Field Studies
of Salt Scaling in Fly Ash Concrete,” Frost Resistance o f
Concrete , M. J. Setzer and R. Auberg, eds., E&FN Spon,
Essen, Germany.
Thomas, M. D. A., 2001 , “Delayed Ettringite Formation
in Concrete: Recent Developments and Future Directions,”
Materials Science o f Concrete , American Ceramics Society,
Westerville, OH.
Thomas, M. D. A., 2006a, “The Role of Calcium in AlkaliSilica Reaction,” Materials Science o f Concrete—The
Sidney Diamond Symposium , American Ceramics Society,
Westerville, OH, pp. 325-331 .
Thomas, M. D. A., 2006b, “The Role of Calcium
Hydroxide in Alkali Recycling in Concrete,” Materials
Science o f Concrete Special Volume on Calcium Hydroxide
in Concrete , J. Skalny, J. Gebauer, and I. Odler, eds., Amer-
ican Ceramic Society, Westerville, OH, pp. 269-280.
Thomas, M. D. A., 201 1 , “The Effect of Supplementary
Cementing Materials on Alkali-Silica Reaction: A Review,”
Cement and Concrete Research , V. 41 , No. 3, Mar., pp.
209-21 6. doi: 1 0.1 01 6/j.cemconres.2011 .03.001
Thomas, M. D. A., and Bleszynski, R. F., 2001 , “The Use
of Silica Fume to Control Expansion Due to Alkali-Aggregate Reactivity in Concrete—A Review,” Materials Science
o f Concrete VI, J. Skalny and S. Mindess, eds., American
Ceramics Society, Westerville, OH, pp. 377-434.
Thomas, M. D. A., and Folliard, K. J., 2007, “Concrete
Aggregates and the Durability of Concrete,” Durability o f
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Page and M.from
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Page, eds., Woodhead, Cambridge, UK, pp. 247-281 .
Thomas, M. D. A., and Innis, F. A., 1 998, “Use of the
Accelerated Mortar Bar Test for Evaluating the E ffcacy
of Mineral Admixtures for Controlling Expansion due to
Alkali-Silica Reaction,” Cement, Concrete and Aggregates ,
V. 21 , No. 2, pp. 1 57-1 64.
Thomas, M. D. A.; Hooton, R. D.; Scott, A.; and Zibara,
H., 201 2, “The Effect of Supplementary Cementitious
Materials on Chloride Binding in hardened Cement Paste,”
Cement and Concrete Research , V. 42, No. 1 , Jan., pp. 1 -7.
doi: 1 0.1 01 6/j.cemconres.2011 .01 .001
Thomas, M. D. A., and Skalny, J., 2006, “Chemical Resistance of Concrete,” Signifcance o f Tests and Properties o f
Concrete and Concrete-Making Materials , STP 1 69D, J. F.
Lamond and J. H. Pielert, eds., ASTM International, West
Conshohocken, PA, pp. 253-273.
Thomas, M. D. A.; Fournier, B.; and Folliard, K. J., 2008a,
“Report on Determining the Reactivity of Concrete Aggregates and Selecting Appropriate Measures for Preventing
Deleterious Expansion in New Concrete Construction,” Federal Highways Administration, Report FHWAHIF-09-001 , National Research Council, Washington, DC.
Thomas, M. D. A.; Folliard, K.; Drimalas, T.; and Ramlochan, T., 2008b, “Diagnosing Delayed Ettringite For mation in
Concrete Structures,” Cement and Concrete Research , V. 38,
No. 6, pp. 841 -847. doi: 1 0.1 01 6/j.cemconres.2008.01 .003
Thomas, M. D. A.; Fournier, B.; Folliard, K. J.; Ideker,
J. H.; and Shehata, M., 2006, “Test Methods for Evaluating
Preventive Measures for Controlling Expansion Due to
Alkali-Silica Reaction in Co ncrete,” Cement and Concrete
Research , V. 36, No. 1 0, Oct., pp. 1 842-1 856. doi: 1 0.1 01 6/j.
cemconres.2006.01 .01 4
Thomas, M. D. A.; Fournier, B.; Folliard, K. J.; Shehata,
M. H.; Ideker, J. H.; and Rogers, C., 2007, “Performance
Limits for Evaluating Supplementary Cementing Materials
Using Accelerated Mortar Bar Test,” ACI Materials Journal ,
V. 1 04, No. 2, Mar.-Apr., pp. 1 1 5-1 22.
Thomas, M. D. A.; Hooton, R. D.; and Rogers, C. A., 1 997,
“Prevention of Damage Due to Alkali-Aggregate Reaction
(AAR) in Concrete Construction—Canadian Approach,”
Cement, Concrete and Aggregates , V. 1 9, No. 1 , June, pp.
26-30. doi: 1 0.1 520/CCA1 001 8J
Thomas, M. D. A.; Nixon, P. J.; and Pettifer, K., 1 991 ,
“The Effect of PFA on Alkali-Silica Reaction,” Second
CANMET/ACI Con ference on the Durability o f Concrete ,
SP-1 26, V. M. Malhotra, ed., V. II, American Concrete Institute, Farmington Hills, MI, pp. 91 9-940.
Thornton Jr., H. T., 1 978, “Acid Attack of Concrete
Caused by Sulfur Bacteria Action,” ACI Journal Proceedings , V. 75, No. 1 1 , Nov., pp. 577-584.
Thorvaldson, T., 1 954, “Chemical Aspects of the Durability of Cement Products,” Proceedings, Third Interna tional Symposium on the Chemistry o f Cement, Cement and
Concrete Association, London, pp. 436-466.
Thorvaldson, T.; Lamour, R. K.; and Vigfusson, V. A.,
1 927, “The Expansion of Portland Cement Mortar Bars
During Disintegration in Sulphate Solution,” English
Standard
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1 99-206.
Thorvaldson, T.; Wolochow, D.; and Vigfusson, V. A.,
1 929, “Studies on the Action of Sulphates on Portland
Cement: I. The Use of the Expansion Method in the Study
of the Action of Sulphates on Portland Cement Mortar and
Concrete,” Canadian Journal o f Research , V. 1 , No. 3, pp.
273-284. doi: 1 0.1 1 39/cjr29-01 7
Thurmann, M. T., 1 969, Pavement Wear Caused by
Studded Tires (Piggdekkenes Slitasje på Vegdekker) , Norwegian State Highway Laboratory, Oslo, Norway.
Torii, K., and Kawamura, M., 1 994, “Mechanical and
Durability-Related Properties of High-Strength Concrete
Containing Silica Fume,” High-Performance Concrete,
Proceedings o f the Second ACI International Con ference ,
SP-1 49, V. M. Malhotra, ed., American Concrete Institute,
Farmington Hills, MI, pp. 461 -474.
Touma, W. E.; Fowler, D. W.; and Carrasquillo, R. L.,
2001 , “Alkali-Silica Reaction in Portland Cement Concrete:
Testing Methods and Mitigation Alternatives,” Report ICAR
301-1F, International Center for Aggregates Research,
Austin, TX, 520 pp.
Transportation Research Board, 2006, “Control of
Cracking of Concrete, State of the Art,” Transportation
Research Circular , E-C1 07, Transportation Research Board,
Washington, DC, 46 pp.
Trej o, D. , and Pillai, R. , 2004, “Accelerated Chloride
Threshold Testing: Part II – Corrosion Resistant Reinforcement,” ACI Materials Journal , V. 1 01 , No. 1 , Jan. Feb. , pp. 57-64.
GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
Trejo, D., and Reinschmidt, K., 2007, “Justifying
Material Selection for Reinforced Concrete Structures. I: Sensitivity Analysis,” Journal o f Bridge Engi neering , V. 1 2, No. 1 , Jan.-Feb., pp. 31 -37. doi: 1 0.1 061 /
(ASCE)1 084-0702(2007)1 2:1 (31 )
Tremblay, C.; Bérubé, M. A.; Fournier, B.; Thomas, M.
D. A.; and Folliard, K. J., 2007, “Effectiveness of LithiumBased Products in Concrete Made with Canadian Natural
Aggregates Susceptible to Alkali-Silica Reactivity,” ACI
Materials Journal , V. 1 04, No. 2, Mar.-Apr., pp. 1 95-205.
Tremblay, C.; Bérubé, M. A.; Fournier, B.; Thomas, M. D.
A.; and Folliard, K. J., 201 0, “Experimental Investigation of
the Mechanisms by which LiNO 3 is Effective against ASR,”
Cement and Concrete Research , V. 40, No. 4, pp. 583-597.
doi: 1 0.1 01 6/j.cemconres.2009.09.022
Truc, O.; Ollivier, J.-P.; and Nilsson, L.-O., 2000, “Numerical Simulation of Multi-Species Transport through Saturated Concrete during a Migration Test — MsDi ff Code,”
Cement and Concrete Research , V. 30, No. 1 0, Oct., pp.
1 581 -1 592. doi: 1 0.1 01 6/S0008-8846(00)00305-7
Turk, K., and Karatas, M., 201 1 , “Abrasion Resistance
and Mechanical Properties of Self-Compacting Concrete
with Different Dosages of Fly Ash/Silica Fume,” Indian
Journal o f Engineering and Materials Sciences , V. 1 8, No.
1 , pp. 49-60.
Tutti, K., 1 982, “Corrosion of Steel in Concrete,” S-100
44 , Swedish Cement and Concrete Research Institute, Stockholm, Sweden.
Uchikawa, H.; Uchida, S.; and Hanehara, S., 1 989, “Relationship between Structure and Penetrability of Na Ion in
Hardened Blended Cement Paste Mortar and Concrete,”
Proceedings o f the 8th International Con ference on AlkaliAggregate Reaction , K. Okada, S. Nishibayashi, and M.
Kawamura, eds., The Society of Materials Science, Kyoto,
Japan, pp. 1 21 -1 28.
U.S. Bureau of Reclamation (USBR), 1 963, Concrete
Manual: A Manual for the Control o f Concrete Construc tion , seventh edition, U.S. Department of the Interior,
Denver, CO., pp. 1 2-1 3.
U.S. Bureau of Reclamation (USBR), 1 975, Concrete
Manual , eighth edition, U.S. Department of the Interior,
Denver, CO, 627 pp.
U.S. Bureau of Reclamation (USBR), 1 997, Guide to
Concrete Repair , U.S. Department of the Interior, Washington, DC, 99 pp.
Valenza II, J. J., and Scherer, G. W., 2007, “A Review
of Salt Scaling: I. Phenomenology,” Cement and Concrete
Research , V. 37, No. 7, July, pp. 1 007-1 021 . doi: 1 0.1 01 6/j.
cemconres.2007.03.005
van Aardt, J. H. P., and Visser, S., 1 977, “Calcium
Hydroxide Attack of Feldspars and Clays: Possible Relevance to Cement-Aggregate Reactions,” Cement and
Concrete Research , V. 7, No. 6, Nov., Nov., pp. 643-648.
Van Daveer, J. R., and Sheret, G. D., 1 975, “Concrete
Cover Study,” Final Report No. FHWA-DP-15 , Federal
Highway Administration, Washington, DC.
Van Til, C. J.; Carr, B. J.; and Vallerga, B. A., 1 976,
“Waterproof Membranes for Protection of Concrete Bridge
81
Deck—Laboratory Phase,” NCHRP Report No. 165 , Transportation Research Board, Washington, DC, 70 pp.
Vanden Bosch, V. D., 1 980, “Performance of Mortar
Specimens in Chemical and Accelerated Marine Exposure,”
Performance o f Concrete in Marine Environment , SP-65, V.
M. Malhotra, ed., American Concrete Institute, Farmington
Hills, MI, pp. 487-507.
Verbeck, G. J., 1 958, “Carbonation of Hydrated Portland
Cement,” Cement and Concrete , STP-205, ASTM International, West Conshohocken, PA, pp. 1 7-36.
Verbeck, G. J., 1 968, “Field and Laboratory Studies of the
Sulphate Resistance of Concrete,” Performance o fConcrete—
Resistance o f Concrete to Sulphate and Other Environmental
Conditions , Thorvaldson Symposium, University of Toronto
Press, Toronto, ON, Canada, pp. 11 3-1 24.
Verbeck, G. J., 1 975, “Mechanisms of Corrosion of Steel
in Concrete,” Corrosion o fMetals in Concrete , SP-49, American Concrete Institute, Farmington Hills, MI, pp. 21 -38.
Verbeck, G. J., and Helmuth, R. A., 1 969, “Structures and
Physical Properties of Cement Paste,” Proceedings o f the
Fifth International Symposium on the Chemistry o f Cement ,
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Villani, C.; Nantung, T.; and Weiss, W. J., 201 4b, “The
Inf uence o f Deicing Salt Exposure on the Gas Transport in
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Villani, C.; Spragg, R.; Pour-Ghaz, M.; and Weiss, W.
J., 201 4a, “The Inf uence o f Pore Solution Properties on
Drying in Cementitious Materials,” Journal o f the American Ceramic Society , V. 97, No. 2, Feb., pp. 386-393. doi:
1 0.11 11 /jace.1 2604
Vitousova, L., 1 991 , “Concrete Sleepers in CSD Tracks,”
International Symposium on Precast Concrete Sleepers ,
Madrid, Servicio de Publicaciones del Colegio de Ingenieros
de Caminos, Canales y Puertos, pp. 253-264.
Walder, J. S., and Hallet, B., 1 985, “A Theoretical Model
of the Fracture of Rock during Freezing,” Bulletin o f the
Geological Society o f America , V. 96, No. 3, pp. 336-346.
doi: 1 0.11 30/001 6-7606(1 985)96<336:ATMOTF>2.0.CO;2
Wang, H., and Gillott, J. E., 1 991 , “Mechanism of AlkaliSilica Reaction and the Signifcance o f Calcium Hydroxide,”
Cement and Concrete Research , V. 21 , No. 4, July, pp.
647-654. doi: 1 0.1 01 6/0008-8846(91 )9011 5-X
Warren, C. J., and Reardon, E. J., 1 994, “The Solubility of Ettringite at 25°C,” Cement and Concrete
Research , V. 24, No. 8, Aug., pp. 1 51 5-1 524. doi:
1 0.1 01 6/0008-8846(94)901 66-X
Way, S. J., and Cole, W. F., 1 982, “Calcium Hydroxide
Attack on Rocks,” Cement and Concrete Research , V. 1 2, No.
5, Sept., pp. 61 1 -61 7. doi: 1 0.1 01 6/0008-8846(82)90022-9
Wehner, B., 1 966, “Beanspruchung der Strassenooer f aech
durch Winterreifen mit Spikes,” Technische Universitaet
Berlin, Institute fur Strassen und Verkehrswesen, Germany.
Weyers, R. E.; Pyc, W.; Sprinkel, M. M.; and Kirkpatrick, T. J., 2003, “Bridge Deck Cover Depth Specifcations,”
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GUIDE TO DURABLE CONCRETE (ACI 201.2R-1 6)
Whiting, D. A., and Nagi, M. A., 1 998, Manual on Control
Portland Cement Association, Skokie, IL, 48 pp.
Whiting, D. A., and Stark, D. C., 1 983, “Control of Air
Content in Concrete,” NCHRP Report No. 258 , Transportation Research Board, National Research Council, Washington, DC, May.
Wiebenga, J. G., 1 984, “Durability of 64 Concrete
Constructions On-Shore and Off-Shore,” Cement (s-Herto genbosch) , V. 36, No. 4, Apr., pp. 21 5-21 8. (in Dutch)
Wiens, U.; Breit, W.; and Schiessl, P., 1 995, “Inf uence
of High Silica Fume and High Fly Ash Contents on Alkalinity of Pore Solution and Protection of Steel against Corrosion,” Fifth CANMET/ACI International Con ference on Fly
Ash, Silica Fume, Slag and Natural Pozzolans in Concrete ,
SP-1 53, V. M. Malhotra, ed., V. 2, American Concrete Institute, Farmington Hills, MI, pp. 741 -761 .
Wig, R. J.; Williams, G. M.; and Finn, A. N., 1 91 7, “Durability of Cement Draintile and Concrete in Alkali Soils,”
Technologic Papers o f the Bureau o f Standards No. 95, US
Department of Commerce, Washington, DC, 1 40 pp.
Wilk, W., 1 978, “Consideration of the Question of
Skid Resistance of Carriageway Surfaces, Particularly
of Concrete,” Betonstrassen AG, No. 117, Monograph,
Wildegg, Switzerland.
Williams, G. M., and Furlong, I., 1 926, “Durability of
Cement Drain Tile and Concrete in Alkali Soils,” Fourth
o f Air Content in Concrete , EB11 6,
Progress Report, Technologic Papers o f the Bureau o f Stan Get moreofFREE
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dards No. 307, U.S. Department
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Williamson, G. S.; Weyers, R. E.; Sprinkel, M. M.; and
Brown, M. C., 2003, “Concrete and Steel Type Inf uence
on Probabilistic Corrosion Service Life,” ACI Materials
Journal, V. 1 06, No. 1 , Jan.-Feb., pp. 82-88.
Willis, T. F., and Lord, G. W., 1 951 , “Calculation of Air
Bubble Size Distribution from Results of Rosiwal Traverse
of Aerated Concrete,” ASTM Bulletin 177, Philadelphia, PA,
pp. 56-61 .
Winkler, E. M., 1 997, Stone in Architecture: Properties
and Durability , Springer, New York, 31 3 pp.
Witte, L. P., and Backstrom, J. E., 1 951 , “Some Properties Affecting the Abrasion Resistance of Air-Entrained
Concrete,” ASTM Proceedings , V. 51 , pp. 1 1 41 -11 55.
Wolsiefer Sr., J. T., 1 991 , “Silica Fume Concrete: A Solution
to Steel Reinforcement Corrosion in Concrete,” Durability o f
Concrete , Proceedings of the Second CANMET/ACI International Conference, SP-1 26, V. M. Malhotra, ed., American
Concrete Institute, Farmington Hills, MI, pp. 527-558.
Woods, H., 1 968, Durability o f Concrete Construction,
Monograph No. 4, Iowa State University Press, Ames, IA,
1 87 pp.
Wu, H.; Huang, B.; Shu, X.; and Dong, Q., 201 1 , “Laboratory Evaluation of Abrasion Resistance of Portland
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Engineering , V. 23, No. 5, May, pp. 697-702. doi: 1 0.1 061 /
(ASCE)MT.1 943-5533.000021 0
Xu, Z., and Hooton, R. D., 1 993, “Migration of Alkalki
Ions in Mortar Due to Several Mechanisms,” Cement and
Concrete Research , V. 23, No. 4, Apr., pp. 951 -961 . doi:
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Con ference on Alkali-Aggregate Reactivity , M. A. Bérubé,
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Xu, Z. ; Shen, Y. ; Lu, D. ; Deng, M. ; Lan, X. ; Hu, R. ; and
Tang, M. , 1 998, “Investigation on A New Test Method for
Determining the Alkali Silica Reactivity of Aggregates,”
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V. 20, No. 2, pp. 1 -7.
Yen, B. C., and Bright, R. E., 1 990, “Residential Foundation Deterioration Study for the Cities of Lakewood, La
Palma, and Cypress, California,” California State University, Long Beach, CA, Apr., 1 11 pp.
Yen, T.; Hsu, T.; Liu, Y.; and Chen, S., 2007, “Inf uence
of Class F Fly Ash on the Abrasion-Erosion Resistance
of High-Strength Concrete,” Construction & Building
Materials , V. 21 , No. 2, Feb., pp. 458-463. doi: 1 0.1 01 6/j.
Standard
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conbuildmat.2005.06.051
Yilmaz, V. T., and Glasser, F. P., 1 990, “Reaction of
Alkali-Resistant Glass Fibres with Cement. Part 2. Durability in Cement Matrices Conditioned with Silica Fume,”
Glass Technology , V. 32, No. 4, pp. 1 38-1 47.
Yingling, J.; Mullings, G. M.; and Gaynor, R. D., 1 992,
“Loss of Air Content in Pumped Concrete,” Concrete International, V. 1 4, No. 1 0, Oct., pp. 57-61 .
Young, F. R., 1 989, Cavitation , McGraw-Hill, London,
pp. 45-56.
Zhang, J., and Taylor, P., 201 2, “Investigation of the Effect
of the Interfacial Zone on Joint Deterioration of Concrete
Pavements,” Proceedings o fthe International Con ference on
Long-Li fe Concrete Pavements , Federal Highways Administration, Washington, DC.
Zhang, M.-H., and Gjørv, O., 1 991 , “Effect of Silica
Fume on Pore Structure and Chloride Diffusivity of
Low-Porosity Cement Pastes,” Cement and Concrete
Research , V. 21 , No. 6, Nov.-Dec., pp. 1 006-1 01 4. doi:
1 0.1 01 6/0008-8846(91 )90060-U
Zivica, V., and Bajza, A., 2001 , “Acid Attack of Cement
Based Materials—A Review: Part I. Principle of Acidic
Attack,” Construction & Building Materials , V. 1 5, No. 8,
Dec., pp. 331 -340. doi: 1 0.1 01 6/S0950-061 8(01 )0001 2-5
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