201.2R-16 · 2020. 5. 12. · ACI 201.2R-16 Guide to Durable Concrete Reported by ACI Committee 201...

Guide to Durable ConcreteReported by ACI Committee 201

AC

I 20

1.2R

-16

First PrintingNovember 2016

ISBN: 978-1-945487-39-2

Guide to Durable Concrete

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This guide describes specific types of concrete deterioration. Each chapter contains a discussion of the mechanisms involved and the recommended requirements for individual components of concrete, quality considerations for concrete mixtures, construction proce-dures, and influences of the exposure environment, which are all important considerations to ensure concrete durability.

This guide was developed for conventional concrete but is gener-ally 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.

CONTENTS

CHAPTER 1—INTRODUCTION AND SCOPE, p. 21.1—Introduction, p. 21.2—Scope, p. 3

CHAPTER 2—DEFINITIONS, p. 32.1—Definitions, p. 3

CHAPTER 3—MASS TRANSPORT, p. 33.1—Introduction, p. 33.2—Transport processes in nonreactive porous media, p. 43.3—Factors affecting mass transport in concrete, p. 53.4—Measurement of transport properties, p. 83.5—Obtaining durable concrete, p. 10

CHAPTER 4—FREEZING AND THAWING OF CONCRETE, p. 10

4.1—Introduction, p. 10

Thomas J. Van Dam, Chair R. Douglas Hooton, Secretary

ACI 201.2R-16

Guide to Durable Concrete

Reported by ACI Committee 201

Reza AhrabliJames M. Aldred

Jon B. ArdahlMohamed Bassuoni

Bruce BlairAndrew J. BoydPaul W. Brown

Ramon L. CarrasquilloRachel J. Detwiler

Jonathan E. Dongell

Thano DrimalasKevin J. Folliard

Harvey H. HaynesJason H. Ideker

Francis InnisDonald J. Janssen

Roy H. KeckMohammad S. KhanKimberly E. KurtisMichael L. Leming

Tyler LeyDarmawan Ludirdja

Mohamad NagiRobert E. Neal

Charles K. NmaiKarthik H. Obla

Robert C. O’NeillKyle Austin RidingDavid A. RothsteinHannah C. Schell

Lawrence L. SutterDavid G. Tepke

Michael D. A. ThomasPaul J. Tikalsky

David TrejoOrville R. Werner II

Terry J. WillemsMichelle L. Wilson

Consulting MembersW. Barry ButlerBernard ErlinOdd E. Gjorv*

William G. Hime

Charles J. HookhamAlexander M. Leshchinsky

Stella Lucie MarusinHoward H. Newlon Jr.

Mauro J. ScaliGeorge V. Teodoru

Niels ThaulowJ. Derle Thorpe

Claude B. Trusty Jr.

*Deceased.

ACI Committee Reports, Guides, and Commentaries are intended for guidance in planning, designing, executing, and inspecting construction. This document is intended for the use of individuals who are competent to evaluate the significance 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.

ACI 201.2R-16 supersedes ACI 201.2R-08 and was adopted and published November 2016

Copyright © 2016, 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.

1

4.2—Frost attack of concrete made with durable aggre-gates, p. 11

4.3—Frost attack of concrete made with nondurable aggregates, p. 17

CHAPTER 5—ALKALI-AGGREGATE REACTION, p. 19

5.1—Introduction, p. 195.2—Types of reactions, p. 195.3—Evaluating aggregates for potential alkali-aggregate

reactivity, p. 225.4—Preventive measures, p. 255.5—Tests for evaluating preventive measures, p. 285.6—Protocols for minimizing the risk of alkali-aggregate

reactivity, p. 29

CHAPTER 6—SULFATE ATTACK, p. 306.1—External sulfate attack, p. 306.2—Internal sulfate attack, p. 366.3—Seawater and brine exposure, p. 37

CHAPTER 7—CHEMICAL ATTACK, p. 397.1—General, p. 397.2—Seawater, p. 397.3—Acid attack, p. 417.4—Fresh water, p. 427.5—Carbonation, p. 427.6—Industrial chemicals, p. 437.7—Deicing and anti-icing chemicals, p. 447.8—Environmental structures, p. 45

CHAPTER 8—PHYSICAL SALT ATTACK, p. 458.1—Introduction, p. 458.2—Occurrence, p. 468.3—Background, p. 478.4—Mechanism, p. 478.5—Recommendations, p. 48

CHAPTER 9—CORROSION OF METALS AND DEGRADATION OF OTHER MATERIALS EMBEDDED IN CONCRETE, p. 48

9.1—Introduction, p. 489.2—General principles of corrosion initiation in concrete,

p. 489.3—Propagation of corrosion, p. 499.4—Corrosion-related properties of concreting materials,

p. 499.5—Mitigating corrosion, p. 509.6—Corrosion of prestressed steel reinforcement, p. 539.7—Degradation of materials other than steel, p. 539.8—Summary, p. 54

CHAPTER 10—ABRASION, p. 5410.1—Introduction, p. 5410.2—Testing concrete for resistance to abrasion, p. 5510.3—Factors affecting abrasion resistance of concrete,

p. 55

10.4—Recommendations for obtaining abrasion-resistant concrete surfaces, p. 57

10.5—Studded tire and tire chain wear on concrete, p. 5810.6—Skid resistance of pavements, p. 5810.7—Erosion, p. 59

CHAPTER 11—SUMMARY, p. 60

CHAPTER 12—REFERENCES, p. 60Authored documents, p. 62

CHAPTER 1—INTRODUCTION AND SCOPE

1.1—IntroductionConcrete 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, finished, and cured concrete is capable of 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 chem-ical 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 work synergistically. Depending on the nature of the attack, distress may be concentrated in the paste, aggregate, or rein-forcing components of the concrete, or a combination thereof.

The various factors influencing 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 specific structural element is to be exposed. 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 require-ments, and expected exposure environments, both macro and micro, should be determined before defining the neces-sary materials and mixture proportions.

The use of good materials and proper mixture proportioning will not, by itself, ensure durable concrete. Appropriate place-ment practices and workmanship are essential to the produc-tion of durable concrete. Fresh concrete can be consolidated and molded to the shape desired to serve its intended purpose. During this stage, a number of properties significantly influ-encing the durability of the hardened concrete are established. Pore structure development, air-void system formation, mate-rial uniformity, and potential for cracking are established at early ages and are important to the ultimate durability of

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2 GUIDE TO DURABLE CONCRETE (ACI 201.2R-16)

concrete. As such, durable concrete requires the application of good quality control during construction. Inspection and testing by trained and certified personnel can help ensure the use of durable mixtures and proper practices.

1.2—ScopeThis guide discusses the important mechanisms of concrete

deterioration and gives recommendations on how to mitigate or minimize such damage. This guide also addresses dura-bility by first discussing the importance of mass transport and then addressing specific modes of attack in separate chapters. These include freezing and thawing, alkali-aggre-gate 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 216.1 and cracking is covered in ACI 224R and ACI 224.1R. While cracking does impact the durability of concrete in severe exposures, the different causes of cracking and their specific impacts are not discussed. Cracking is only mentioned in general terms regarding its impact on fluid ingress.

CHAPTER 2—DEFINITIONS

2.1—DefinitionsACI provides a comprehensive list of definitions through

an online resource, “ACI Concrete Terminology,” https://www.concrete.org/store/productdetail.aspx?ItemID=CT16. Definitions provided herein complement that source.

advective transport––transfer of heat or matter via the bulk motion of a fluid.

alkali loading (or content)––total amount of equivalent alkalis (Na2Oe) in a concrete mixture expressed as mass per volume.

calcium sulfoaluminate cement––product obtained by pulverizing clinker containing mainly ye′elimite [Ca4(AlO2)6SO4] 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 of the bulk motion of a fluid.

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 compo-nents such as calcium hydroxide from concrete.

permeability—the ability of a given concrete to permit liquids or gases to pass through.

permeation––flow of 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

concrete pore solution having a sufficiently high concen-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 pris-matic hexagonal crystals typically as acicular radiating groups, with the chemical formula {[Ca3Si(OH)6·12(H2O)] (SO4)(CO3)}.

CHAPTER 3—MASS TRANSPORT

3.1—IntroductionConcrete is a multiphase porous medium consisting of a

multiscale porous cement paste matrix with aggregate inclu-sions. Liquid and gas may be present in any pores and micro-cracks. As such, it is susceptible to the ingress and move-ment of substances (fluids 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. 1996; Baer 1988; Hearn et al. 2006; Hall and Hoff 2012). 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 of fluids and ions will increase by the presence of cracks. However, the specific influences of different types of 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 hard-ened cement paste and because of osmotic pressures gener-ated as water flows to sites of higher alkalinity (Powers et al. 1954; Powers 1975; Helmuth 1960b,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 trans-ported through the concrete, viscosity of the fluid, valence of the ions, and other ionic species present also affect the trans-port 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.


GUIDE TO DURABLE CONCRETE (ACI 201.2R-16) 3

3.2—Transport processes in nonreactive porous media

This section provides a brief overview of the transport processes of fluids (gases and liquids) and ions in solution within a nonreactive porous medium. This is a simplifying assumption because concrete changes chemically and physi-cally 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 fluids or ions are not discussed herein. The trans-port pathways described here include:

a) Transport by permeationb) Advective transportc) Hydrodynamic dispersiond) Diffusion within the porese) Transport due to electrostatic interactions or electrical

migrationMartín-Pérez et al. (2001) modeled transport related to

corrosion of reinforcement in concrete based on chloride transport, moisture diffusion, heat transfer, and oxygen transport using a two-dimensional finite element model. They used a modified version of Fick’s second 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 elec-trical potential, convection, effects of changes in moisture content, and mass exchange of ions between solution and solid hydration phases.

Chung and Consolazio (2005) developed a finite difference model to simulate heat and mass transport in rapid heating conditions, such as fires in reinforced concrete structures. The model accounts for the interference between liquid and gas phases, slip-flow effects in steam flow, and the interference of steel reinforcement in moisture movement in concrete.

3.2.1 Transport by permeation—Permeation is the flow of a fluid under the action of a pressure gradient. Permeability is the property that characterizes the ease with which fluid passes through a porous material under a pressure gradient. For a steady laminar flow through a saturated porous medium, the fluid flow is related to the hydraulic pressure gradient according to Darcy’s law.

dq/dt = K1A∆h/l (3.2.1)

where dq/dt is the flow (expressed as a rate); K1 is the perme-ability coefficient; A is the cross-sectional area; ∆h is the hydraulic head; and l is the thickness of the specimen. The permeability coefficient K1 is the rate of discharge of water under laminar flow conditions through a unit cross-sectional area of a porous medium under a unit hydraulic gradient and standard temperature conditions. Darcy’s law indicates that for a given cross-sectional area and permeability coeffi-cient, the flow is proportional to the hydraulic gradient ∆h/l. Under service conditions, flow is three-dimensional and the concrete may not be saturated. In concrete, the permeability coefficient may change with increased hydration, cracking,

or changes in the pore structure due to various physical and chemical processes.

Permeability coefficients of plastic portland cement pastes of 0.5 water-cement ratio (w/c), calculated from measurements of bleeding, ranged from 5 to 8 × 10–7 m/s for four cements with different chemical composition but the same specific surface (180 m2/kg by the Wagner turbidimeter). The perme-ability coefficient of mature paste (for example, at greater than 28 days of age) is between 1 millionth and one 10 millionth of that of fresh paste. It ranges from 1 × 10–15 to 1.2 × 10–12 m/s for w/c ranging from 0.3 to 0.7 (Powers et al. 1954).

To understand the effects of microstructure on the perme-ability of concrete, Bentz et al. (1999) at the National Insti-tute 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 1992). This maze consists of areas that can allow free movement, as well as areas that impede the transport of a fluid to different degrees. These models can be made in two and three dimen-sions and can include the effects of cracks. Lu et al. (2012) 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 observe the effects on the ability of a fluid to move through the system.

A graphical representation of a simple percolation theory model has been used by Bentz (2000) to model the transport of a fluid through concrete in two dimen-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 flow. This transport process is related directly to the velocity

Fig. 3.2.1‒Graphical representation of a simple percolation theory model used to model the permeability of concrete.



of fluid flow and concentration of ions in solution. Depending on exposure condition and degree of saturation, the velocity of fluid flow depends on the sorptivity or permeability of concrete, viscosity of the fluid phase, and pressure head of the permeating fluid or rate of evaporation from the exposed surface in the case of wick action (Buenfeld et al. 1995).

3.2.3 Hydrodynamic dispersion—Dispersion is the spreading of the ion concentration during advective transport due to variations in the pore fluid velocity. These variations can be a result of the tortuosity of the pore structure, the connec-tivity of the pore network, or variation in the fluid properties.

3.2.4 Diffusion within pores—Diffusion refers to the trans-port 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/dt = De · d2c/dx2 (3.2.4)

where c is the concentration of the ion at distance x from the surface after time t, and De is the effective diffusion coeffi-cient 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 measure-ments obtained under controlled conditions in the laboratory, diffusion coefficients increase with temperature and water-to-cementitious materials ratio (w/cm) and decrease with increasing degree of hydration. Because concrete pore solu-tions have high ionic strength, electrical charge effects can be significant. Diffusion coefficients 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 experimen-tally using Fick’s second law are generally termed “apparent diffusion coefficients”. Diffusion coefficients for Na+ in concrete are on the order of 10–11 to 10–13 m2/s, and for Cl–, on the order of 10–11 to 10–12 m2/s (Taylor 1997).

3.2.5 Transport due to electrostatic interactions or elec-trical migration—Migration refers to the transport mecha-nism 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-

strated by Snyder (2001) and Snyder and Marchand (2001). Electrical migration occurs when an external electric field such as in ASTM C1202 (or AASHTO T277) is applied to the medium (Buenfeld et al. 1998). The migration flux Ji of ion i is given by

J DC

z FRT xi i ii= −

∂∂ϕ (3.2.5)

where ∂φ/∂x is the potential difference; Zi is the charge of the ion; F is the Faraday constant; T is the temperature; Di is the ion diffusivity; R is the gas constant; and Ci is the concentra-tion of the ion in solution.

Further information on ionic transport in concrete can be found in McGrath and Hooton (1996) and for the more complex case of multi-species transport in Truc et al. (2000) and Samson et al. (1999).

3.3—Factors affecting mass transport in concrete3.3.1 Porosity and pore size distribution—Porosity is

defined as the volume of voids as a fraction that is usually expressed as a percent of the total volume

porosity (%) = (volume of voids/total volume) × 100% (3.3.1)

Figure 3.3.1a 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 most significant effect on the transport properties. Trans-port properties, however, depend more on the connectivity of the pores than on either the porosity or size of the pores. Figure 3.3.1b 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 fill some of the void space formerly occu-pied by water. With time, this process results in a continued

Fig. 3.3.1a–‒Relative sizes of different types of pores and other microstructural features (adapted from Mehta [1986]).



decrease in the porosity of the cement paste. Figure 3.3.2a (Mehta 1986) 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 100 percent to reach the same porosity. The degree of hydration that could be expected for good curing conditions—for example, moist curing for 5 to 7 days—would range between 70 and 80 percent, depending on cement chemistry and fineness, and on hydration temperatures; a 100 percent degree of hydration is not a practical possibility. Figure 3.3.2b shows the relationship between porosity and permeability. Above a porosity of approximately 30 percent, the coefficient of permeability increases sharply.

Powers (1962b) calculated that for cement paste with a w/c of 0.38, all of the capillary pore space was just filled 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 proportion-ately less gel and more capillary space. Powers et al. (1959) 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 temperature3.3.3.1 At normal temperatures—Soon after mixing

cement with water, a gel layer forms on the surfaces of the cement grains (Taylor 1997). 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 CH is accompanied by significant evolution of heat. The CH forms massive crystals in the originally water-filled space. The C-S-H forms a thickening layer around

the cement grains. As the shells grow outward, they begin to coalesce at about 12 hours, a time coinciding with the maximum rate of heat evolution (Fig. 3.3.3.1) and corre-sponding approximately to completion of setting (Taylor 1997). In Fig. 3.3.3.1, the first heat peak is associated with the initial hydrolysis of the C3S and the hydration of the C3A. The acceleration period begins with the renewed evolution of heat at the beginning of the second peak as the initial hydration products of the C3S begin to form. Initial

Table 3.3.2–Approximate age required to produce maturity at which capillaries become discontinuous for concrete continuously moist-cured (Powers et al. 1959)

w/c by mass Time required

0.40 3 days

0.45 7 days

0.50 14 days

0.60 6 months

0.70 1 year

Over 0.70 Impossible

Fig. 3.3.2b‒–Both compressive strength and permeability are related to the capillary porosity of the cement paste (adapted from Powers [1958]).

Fig. 3.3.2a‒–Water-cement ratio versus capillary porosity for cement paste at different degrees of hydration (Mehta 1986) based on equations developed by Powers and Brown-yard (1948).

Fig. 3.3.1b–‒Porosity and permeability are related but distinct. The two hypothetical materials shown have approxi-mately the same porosity (total volume of pores), but different permeabilities. Discrete pores, such as those resulting from air entrainment, have almost no effect on permeability, but interconnected pores increase permeability.



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 associ-ated with the renewed formation of ettringite (Taylor 1997). Further hydration of the cement continues at a much slower rate, asymptotically approaching 100 percent (Mindess and Young 1981). After the spaces between the hydration shells and the cement grains fill with hydration products, further hydration is slow (Taylor 1997). Capillary pores remaining in mature cement paste increase in size with w/c and have diameters of 10 nm and higher (Mindess and Young 1981).

3.3.3.2 At high or low temperatures—Like most chem-ical reactions, cement hydration is faster with increasing temperature. Verbeck and Helmuth (1969) postulated that at elevated temperatures, cement hydration products would not have time to diffuse any significant distance 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 (1981) found that the total porosities of pastes hydrated at 140°F (60°C) were greater than those of pastes hydrated at 81°F (27°C). For cement pastes hydrated at low tempera-tures, on the order of 50°F (10°C), the hydration products are more evenly distributed and the pores fine and discon-tinuous. For cement pastes hydrated at elevated tempera-tures, pores are coarser and more interconnected (Kjellsen et al. 1991). Cement pastes containing fly ash, slag cement, or both, are less sensitive to the effects of elevated tempera-tures, as discussed in 3.3.5.

3.3.4 Aggregates—Aggregates generally have fundamen-tally 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 (1986) 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

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 of water-filled space beneath aggregate particles; less efficient packing of particles of 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. 1995). As the cementitious materials hydrate, the interfacial transi-tion zone fills preferentially with CH and ettringite. Because 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 fluids 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 aggre-gate and cement paste (Mehta 1986). It was found that the diffusivity of chloride in the interfacial transition zone is 10 times greater than that in bulk cement paste (Delagrave et al. 1997). The connectivity of pores in the interfacial transition zone may be high, leading to significantly greater rates of transport for some concretes than might be predicted from their mixture proportions. For mixtures with high coarse

Fig. 3.3.3.1––Heat evolution of Type I/II portland cement paste as measured by conduction calorimetry (Image cour-tesy of E. Shkolnik.).

Fig. 3.3.4‒–Representation of transition zone at paste/aggregate interface in concrete, showing more coarsely crystalline and porous microstructure than in interzonal mass (Mehta 1986).



aggregate contents, such as paving concretes, flow through the interfacial transition zone can completely dominate moisture movement in the concrete over a wide range of w/c (Janssen and Snyder 1994).

3.3.5 Supplementary cementitious materials—The use of SCMs can significantly reduce the permeability and diffu-sivity of concrete. These materials may not reduce the total porosity to any great extent, but instead act to refine and subdivide the pores so that they become less continuous. Some benefits are obtained by the improvements in work-ability, particularly the reduction in bleeding, afforded by these materials. The use of fly ash can also reduce the water demand of some concrete mixtures, allowing reductions in water-to-cementitious materials ratio (w/cm) while main-taining equivalent workability. The lower specific gravity of SCMs relative to cement results in an increased volume of solids for a given w/cm with increasing dosage. The greatest benefits from the standpoint of 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 results in a finer system of capillary pores.

Slag cement and alumina-bearing pozzolans have an addi-tional benefit as the hydration products are highly effective in binding chloride ions, preventing further penetration into the concrete (Thomas et al. 2012). This effect is particularly important in applications such as parking garages, bridge decks, and marine construction, where the reinforcing steel is vulnerable to chloride-induced corrosion.

The indiscriminate use of SCMs is not necessarily bene-ficial. Their pozzolanic and hydraulic reactions take time, partly because the cement must hydrate sufficiently to produce CH to participate in the reaction, and partly because SCMs may vary significantly in terms of kinetics or rate of reactivity. The engineer must consider the properties of the concrete at early ages. Because the pozzolanic reac-tions 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 (1995) found that both silica fume and slag were effective in refining the pore structure. Campbell and Detwiler (1993) 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 magni-tude, with CSA Type 10 (now designated Type GU) portland cement alone performing the worst, and optimized blends of cement, slag, and silica fume performing the best. They did not use fly ash in their study.

Bentz et al. (1995) 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 of transport propertiesMeasurement of the transport properties of concrete is

complicated by the interactions between the concrete and the substance that is moving through it; the changing prop-erties of concrete with time; and the sensitivity of transport properties to variations in moisture, temperature, and other conditions. Because many tests accelerate the transport mechanism of the fluid 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 assumption of laminar flow used in many calculations (Eq. (3.2.1)). Furthermore, laboratory tests are often conducted under highly controlled conditions that may not accurately reflect actual service conditions. When used 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 construc-tion (Puerto Rico DOT SP934). The following sections discuss available ASTM tests used to characterize transport properties, along with commonly used variations.

3.4.1 Ions3.4.1.1 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 C1202 (or AASHTO T277), in which a cylindrical specimen with a nominal diameter of 4 in. (100 mm) and a length of 2 in. (50 mm) is vacuum satu-rated 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 indi-cation of the chloride ion penetrability of the concrete.

Essentially, ASTM C1202 (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 measure-ment. While ion diffusion depends primarily on the micro-structure and chemical binding capacity of the matrix, elec-trical conductivity depends on both microstructure and pore solution chemistry (Buenfeld and Newman 1987). Different proportions of SCMs can profoundly affect the pore solu-tion chemistry. For example, Page and Vennesland (1983) found that 10 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 (1993) 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 attrib-uted the differences between the two sets of test results to differences in electrical conductivity due to differences in pore solution chemistry.



Corrosion inhibitors such as calcium nitrite add ions to the pore solution that increase the pore fluid 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 micro-structural effects as well as changes in the pore solution chemistry (Ann et al. 2006; Reou and Ann 2008). If ASTM C1202 (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 diffusion; thus, the presence of flaws in the concrete can significantly affect the results of the test. In addition, the measurement is taken before steady-state conditions are reached (Zhang and Gjørv 1991).

Low-quality concrete is difficult to evaluate properly because the temperature of the specimens rises when current is applied, increasing the rate of diffusion. McGrath and Hooton (1999) proposed a modification to reduce the test period to 30 minutes to eliminate this problem. High-quality concretes may be difficult to distinguish from one another because the total charge passed is so low, and because the test results are variable (Hooton 1989).

ASTM C1202 (or AASHTO T277) has been criticized for many of the previously-cited interferences (Andrade 1993; Feldman et al. 1994; Streicher and Alexander 1994; Shi 2004).

Despite its limitations, ASTM C1202 (or AASHTO T277) is rapid, convenient, and according to Hooton (1989), what-ever property it is measuring probably is coincident with permeability. If properly interpreted, it can be used effec-tively for quality control in construction (Bognacki et al 2010), although it should not be used to make fine distinc-tions among concretes and cannot be used to compare concretes made with different materials or mixture propor-tions, 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 C1543 (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 environ-ment 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 (12 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 C1556, in which a core from the slab is milled or sliced to obtain samples at eight depths for the purpose of determining the apparent chloride diffusion coeffi-

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 chlo-ride 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 C1543 is its duration. As the specimens must be cured for 14 days and then dried for 28 days before the beginning of the ponding, the 90-day version of the test takes 118 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 180 days.

Although the ponding test does provide a crude one-dimensional profile of chloride ion ingress, the profile is not a reflection of chloride diffusion alone. The initial mode of 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. 1995). Diffusion of the chlorides also takes place. McGrath and Hooton (1999) 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 significantly increases the apparent diffusion coefficient, especially for concretes containing SCMs (Ngala and Page 1997). For high-quality concretes, it may be difficult to develop a chloride profile based on a 90-day ponding period because so little chloride penetrates into the concrete. Extending the ponding period to 180 days and increasing the number of samples taken help to resolve this problem (Berke and Hicks 1992; Andrade and Whiting 1996; Sherman et al. 1996; McGrath and Hooton 1999).

Precision of the sampling can make a significant differ-ence in the conclusions drawn from the results. In analyzing their data from AASHTO T259 (a predecessor of ASTM C1543), McGrath and Hooton (1999) showed that imprecise sampling makes it difficult to distinguish between a high-quality 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 specified in ASTM C1556 avoids this problem.

ASTM C1556 avoids some of the problems associated with ASTM C1543, as the concrete specimen is sealed on all sides except the finished surface, and is saturated with a CH solution before exposure to the sodium chloride solu-tion. Thus, the chlorides penetrate into the specimen only by diffusion, not by sorption. The specimen is then placed in a concentrated (165 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-perfor-mance concretes containing SCMs. In practice, the exposure time could be extended to 90 days or longer.

The sampling is more precise than for ASTM C1543, as indicated previously. A nominal 4 in. (100 mm) diam-



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 diffusion coefficient is determined using nonlinear regression analysis to fit the data to Fick’s second law.

The use of Fick’s second law to reduce the data is a conve-nient but questionable practice. Pettersson (1994) noted that the applicability of Fick’s second law, which is a simplifi-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 perme-able 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 diffu-sant occurs.

Pettersson (1994) further noted that all three of these assumptions are violated in the diffusion of chloride ions through concrete. That is, concrete is heterogeneous, its diffu-sion properties change with time and with the concentration of the diffusant, and both chemical reactions and physical binding can occur. ASTM C1556 uses the term “apparent chloride diffusion coefficient” to make clear that the result obtained is not a true diffusion coefficient. Although ASTM C1556 in some cases may be conducted in less time than ASTM C1543, it is still a lengthy test.

3.4.2 Fluids— One test for the absorption of water by hard-ened concrete is ASTM C642, in which a piece 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 deter-mined. 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 C1585 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. (100 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. (1999) 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 assumes laminar flow. This simplifying assumption is not strictly true because in very small pores the flow is partly molecular in nature. Gas permeability varies with the degree of saturation of the concrete; the coefficient of 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 concreteObtaining durable concrete for given conditions of expo-

sure requires suitable mass transport properties for the fluids and ions of interest. Proper attention to all aspects of good concrete practice is important. The most significant factors, however, are an appropriately low w/cm; judicious use of SCMs; and good workmanship, including mixing, place-ment, compaction, and curing. Elevated curing temperatures can be deleterious to the transport properties of concrete, although the use of an appropriate combination of cementi-tious 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 condi-tions. 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 condi-tions, and appropriate structural design and detailing can minimize the width of cracks. Further guidance is avail-able from the Transportation Research Board (2006) and Detwiler and Taylor (2003).

CHAPTER 4—FREEZING AND THAWING OF CONCRETE

4.1—IntroductionDeterioration of concrete exposed to freezing can occur

when there is sufficient internal moisture 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 thermody-namic conditions to provide a sufficient degree of saturation 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 sufficient maturity before being exposed to freezing, such as concrete for columns or floor 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 severity of exposure should be quantified by a combination of freezing, which is the number of annual cycles of freezing plus average low temperature reached



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 protec-tion from steps taken to limit the fractional volume of freez-able water in the concrete. A description of damage due to freezing and thawing of concrete made with durable aggre-gates, 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 of aggregate identification procedures, is given in 4.3.

4.2—Frost attack of concrete made with durable aggregates

4.2.1 Description of frost damage4.2.1.1 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 concrete4.2.1.2.1 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 considerably accelerated by deicing salts. Vehicle traffic 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 before the first exposure to freezing temperatures.

4.2.2 Preventing frost damage in new concrete4.2.2.1 Protection from early freezing—Young concrete

should be protected from freezing by following the proce-dures and maintaining the minimum temperatures recom-

mended in ACI 306R. After consolidation and finishing, 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 before freezing—Adequate curing, including preventing excessive drying and main-taining adequate temperature, will ensure that the concrete has hydrated sufficiently to substantially 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 1962a). 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 antic-ipated, 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 19.3.2.1 in ACI 318-14 refers to the specified design strength. The two strengths are not the same, as concrete could be exposed to repeated cycles of freezing and thawing at an age significantly earlier than the age associated with the specified design strength.

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 1990). 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 expo-sure and 0.45 for severe and very severe exposure (Table 4.2.3.1a), 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 318-14).

Limiting the w/cm to a specified maximum has the effect 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 sufficient 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, addi-tives, and supplementary cementitious materials (SCMs)



that permit reduction of the amount of freezable water in concrete. Slag cement or pozzolans can refine 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 beneficial effect of 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

water. Numerous field observations have indicated when surfaces are finished by hand, as opposed to machine-finished or cast against formwork, and exposed to deicing salts (Exposure Class F3a in Table 4.2.3.1b), limits on the SCMs may be required due to a variety of factors that appear to include modification of the entrained air-void system (4.2.3.2) as well as superficial changes to the w/cm. Table 4.2.3.1c gives recommended cementitious materials limita-tions for Exposure Class F3a (Thomas 1997).

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 (1993) and Pigeon (1994) found that the cost and difficulty in placing and finishing 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 of sufficient 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 specific parameters normally used to evaluate an air-void 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

Table 4.2.3.1c—Cementitious materials limitations for Exposure Class F3b

Cementitious materialsMaximum percent of total

cementitious materials by mass*

Fly ash or other pozzolans conforming to ASTM C618 25

Slag conforming to ASTM C989/C989M 50

Silica fume conforming to ASTM C1240 10

Total of fly ash or other pozzolans, slag, and silica fume 50†

Total of fly ash or other pozzolans and silica fume 35†

*The total cementitious materials also include ASTM C150/C150M, ASTM C595/C595M, ASTM C845/C485M, and ASTM C1157/C1157M cements.The maximum percentage should include:(a) Fly ash or other pozzolans in Type IP blended cement, ASTM C595/C595M or ASTM C1157/C1157M(b) Slag used in the manufacture of an IS blended cement, ASTM C595/C595M or ASTM C1157/C1157M(c) Silica fume, ASTM C1240, present in a blended cement†Fly ash or other pozzolans and silica fume shall constitute no more than 25 and 10 percent, respectively, of total mass of the cementitious materials.

Table 4.2.3.1b—Requirements by freezing-and-thawing exposure class

Exposure Class

Minimum fc ,* MPa

(psi)Maximum

w/cm† Air content

Limits on cementitious

materials

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 4.2.3.1c‡

F3b# 32 (4500) 0.45** Table 4.2.3.2.4 No restriction‡

*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 freez-able 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 (mois-ture, temperature, and time) is provided so that the minimum fc is achieved before initial exposure to freezing and thawing.§Hand-finished surfaces.#Formed and machine-finished surfaces.**A lower w/cm may be needed when corrosion is of concern (ACI 318-14).

Table 4.2.3.1a—Freezing-and-thawing exposure classes

Exposure Class Severity Condition

F0 Not applicable Concrete not exposed to freezing conditions

F1 Moderate

Concrete exposed to freezing and thawing conditions, but very low probability of

concrete being near saturation at time of exposure*

F2 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†

F3 Very severe 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 floors in areas protected from direct exposure to moisture.†Examples are: vertical surfaces below the level of snow accumulation; vertical surfaces with sufficient 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.



made for this parameter: the voids are spherical, of equal size, and evenly distributed in a simple cubic lattice throughout the paste (Powers 1949). 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 1954; Backstrom et al. 1954, 1958b). Other researchers (Dubovoy et al. 2002; Attiogbe et al. 1992) have shown that concrete with spacing factors greater than 0.008 in. (0.20 mm) may exhibit satisfactory freezing-and-thawing resistance under specific conditions in laboratory tests. CSA A23.1-14/CSA A23.2 requires an average spacing factor of less than 0.009 in. (0.23 mm) with no single value to exceed 0.010 in. (0.26 mm). This is relaxed to an average value of 0.0098 in. (0.25 mm) for concrete with specified strength greater than 10,000 psi (70 MPa).

4.2.3.2.2 Specific surface α—Specific 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 of specific surface is that all voids are spherical, which makes it a function of average chord length alone (Powers 1949). Specific surface is, there-fore, a good indicator of average void size. As average void size goes up, specific surface goes down. Smaller voids provide compliance with the L requirement at lower air contents. The generally accepted value of specific surface for resistance to freezing and thawing is a minimum of 600 in.2/in.3 (25 mm2/mm3) (Powers 1949), though other researchers have shown that concrete with a specific surface area less than 600 in.2/in.3 (25 mm2/mm3) may exhibit satis-factory freezing-and-thawing resistance in laboratory tests (Dubovoy et al. 2002; Attiogbe et al. 1992).

4.2.3.2.3 Philleo factor F′—Philleo (1955) 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 assump-tion that voids are randomly sized and distributed and estab-lishes 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 (1951) to establish a relationship between air-void chord lengths and voids per unit volume. The equation is used to either deter-mine the percentage of paste that is protected because it is within a specified distance of an air-void, or alternately, to determine the distance from an air-void resulting in a speci-fied percentage of the paste. This distance, called the Philleo factor, is often compared with the spacing factor. In actu-ality, however, the Philleo factor is more sensitive to the true air-void distribution than the air content, paste content, and number of voids that influence 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 difficulty in acquiring the data necessary for calculation. These data consist of a record of all chord

lengths measured in a linear traverse (ASTM C457/C457M). Most petrographers, however, use the alternate modified point count method (ASTM C457/C457M), which does not collect the necessary data. While no specific 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 specific surface values of approximately 600 in.2/in.3 (25 mm2/mm3) 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 1993, 1994).

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 difficult to measure in the field. Total air content is therefore generally specified 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 specified total air content of the concrete mixture is achieved. The actual air content necessary to ensure the production of the necessary air-void system is affected by mixing action, workability of the mixture, cement composition, types and amounts of other admixtures, and others (Whiting and Stark 1983). In addition, concrete handling during transport, placing, and finishing 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 (1952, 1956) 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 specific mixtures, admixtures, and construction proce-dures could warrant other values.

Achieving the total air content specified 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 drip-lines on sidewalks, and structural members directly below unsealed joints of bridge slabs. Be careful to always provide positive drainage of runoff from flat 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.



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 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 prop-erly 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 main-taining 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 1983). Air-entraining 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, stabi-lize, and retain the bubbles.

The system of air bubbles in the fresh concrete at place-ment 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 incorpo-rated 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 of concrete and cleanliness and efficiency of 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. Imme-diately 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. 1958a,b; Backstrom et al. 1958a,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 1990). 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 1990; Dyer 1991; Whiting and Nagi 1998). The possible loss of entrained air can be a major concern for concrete exposed to severe freezing-and-thawing 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, experi-ences 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 form-work. Pumping may also be continuous or interrupted at various times. This sequence of events makes it difficult 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. 1992; Janssen et al. 1995). While changes in the air-void system have been documented in multiple cases (Whiting and Nagi 1998; Janssen et al. 1995; Pleau et al. 1995), evidence of reduced resistance to freezing and thawing has not been identified (Elkey et al. 1993; Whiting and Nagi 1998).

4.2.4.2 Consolidating—On discharge of fresh concrete, air voids could be entrapped in the concrete. While filling forms, it is virtually impossible to avoid the inclusion of air

Table 4.2.3.2.4—Recommended air contents

Nominal maximum aggregate size, in.

(mm)

Air content, percent*

Exposure Class F1Exposure Class F2

and F3

3/8 (9.5) 7 7.5

1/2 (12.5) 7 7

3/4 (19) 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 recom-mendations are based on 18 percent air in the paste portion of the concrete with a Vinsol resin air-entraining agent (from an analysis of work by Klieger [1952]). Mixture proportions based on guidance in ACI 211.1 for angular coarse aggregates along with the maximum w/cm values from Table 4.2.3.1b 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.



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 semifluid mixture can rise to the surface through the temporarily fluid mixture. Second, vibration imposes a cyclic compression in the concrete that locally increases, then decreases, the water pressure surrounding the air bubbles. This causes the bubbles them-selves to compress and decompress at the frequency of the vibrator (Young 1989). 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. (1992) showed that vibration at approximately 6000 vibra-tions 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 summary, vibration of fresh concrete refines the air-void 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 consol-idated, 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 finishing. Repeated passes of a finishing tool can force bubbles together, resulting in fewer, larger bubbles. The air content of the concrete at the surface can be reduced by over-finishing, that is, over-manipulation of the surface can reduce resistance to freezing and thawing. This is particularly the case if the surface is finished while it is still covered with bleed water or if water has been applied to make it easier to finish. Finishing 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 significantly reduced. Thomas (1997) found that this was especially true for hand-finished concrete having a high dosage of SCMs (50 percent Class F fly ash).

Machine-placed and finished areas of the same 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 delami-nate at the surface (ACI 301; ACI 302.1R; Tarr and Farny 2008). Hence, as per ACI 301, concrete for slabs to receive a hard trowel finish should not contain an air-entraining agent or have a total air content greater than 3 percent.

4.2.4.4 Curing—Curing is defined as the maintenance of 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 microstruc-ture 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 suscep-tible 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 conflict with the fact that curing can increase the risk of satu-rating concrete during exposure to early freezing. It is gener-ally 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 (1947) 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 (–18°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 515.2R. Vapor retarders for use under slabs are discussed in ACI 302.1R.

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



moisture that may have entered from unsealed surfaces. An example would be the sealing of the top surface of a slab-on-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-specific conditions.

4.2.5.2 Drainage and other methods—The moisture in concrete can often be sufficiently 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 applica-tion-specific, 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 significantly 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 the wall melts, the remaining snow could 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 subse-quent freezing.

4.2.5.2.3 Redirection of water flow—Downspouts emptying water across a concrete slab can be redirected so that runoff does not flow 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 1966). Historically, the capil-lary-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 1945, 1954, 1955, 1975; Powers and Helmuth 1956; Helmuth 1960a; Litvan 1972; Penttala 1998; Setzer 1999, 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 rela-tively low permeability of cement paste. Distance from

the void boundary, the degree of saturation, and the rate of freezing would influence the magnitude of hydraulic-stress buildup (Cordon 1966; Powers 1945, 1954, 1955).

4.2.6.2 Osmotic pressure—The hydraulic-pressure concept was later modified when experiments showed significant evidence that moisture was moving toward, rather than away from, freezing sites (Powers and Helmuth 1956; Helmuth 1960a; Powers 1975). Marchand et al. (1995) and Penttala (1998) proposed that not all of the water in the capil-lary 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 flow is thought to cause damage. Litvan (1972) 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 concentra-tion gradients.

Penttala (1998), Scherer and Valenza (2005), Setzer (1999, 2002), and Coussy and Monteiro (2008, 2009) have combined the moisture expulsion and osmotic pressure theo-ries 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. Scherer and Valenza (2005) add that when the larger voids fill 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 tempera-tures slightly below 32°F (0°C) may generally dominate the further degradation of previously damaged or microcracked concrete (Litvan 1978). A theoretical model was developed and proven successful for predicting freezing-and-thawing damage in rock (Walder and Hallet 1985; Hallet et al. 1991). 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 migra-tion 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 (1980) study of freezing effects in porous media as applied to soils.

4.2.6.4 Implications of freezing-and-thawing damage mechanisms—Uncertainty and opposing views of funda-mental processes governing freezing in concrete under-mine 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/



C666M, is used to determine the resistance of concrete mixtures to internal damage. ASTM C672/C672M evaluates scaling resistance. The latter relies on visual classifications, sometimes supplemented with mass loss measurements. These tests, like some of their predecessors, are criticized for not adequately representing typical environmental condi-tions. The tests cycle between extreme temperatures too rapidly. The CDF test (RILEM TC 117-FDC 1996) provides greater reproducibility and is more quantitative with respect to scaling measurements. Hallet et al. (1991) proposed that conventional tests give only limited guidance for under-standing the processes governing freezing at field sites 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 field 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 tests and do not perform well in the field, while others fail the tests and perform quite satisfactorily in the field. This is probably because the tests for internal damage and scaling due to freezing and thawing do not address all of the signifi-cant variables. Stark (1989a) 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 of exposure to specified temperature ranges.

4.3—Frost attack of concrete made with nondurable aggregates

Deterioration due to freezing and thawing of properly proportioned, air-entrained concrete made with aggre-gate susceptible to freezing-and-thawing damage is often referred to as D-cracking. Many types of coarse aggregate have been identified as susceptible 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 suscep-tibility 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 of D-cracking4.3.1.1 General description—D-cracking is character-

ized 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

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 flatwork such as highway pave-ments, parking lots, and sidewalks. These are areas that frequently have readily available moisture from precipita-tion 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, mois-ture 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 1987). This is prob-ably 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 some-what 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 character-istic cracks parallel to the joint.

4.3.1.3 Vertical construction—Though much less common, D-cracking can also appear on vertical construc-tion. 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 suscep-tible 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 of D-cracking4.3.2.1 Role of mixture 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 1987; Missouri Highway and Transporta-tion Department 1990). Most coarse aggregates identified as susceptible to D-cracking are sedimentary rocks, although many sedimentary rocks have not been found to be suscep-tible to D-cracking. Igneous rocks are generally not consid-ered 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 identified as susceptible (Stark 1976).

The maximum aggregate size is also important in the development of D-cracking. Numerous studies (Stark and Klieger 1973; Klieger et al. 1974; Stark 1976; Missouri Highway and Transportation Department 1990; Almond and Janssen 1991) have shown that reducing the nominal



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 of aggregate identification—D-cracking can require a number of years to fully develop, during which time much susceptible concrete could be placed before a problem is identified. This, combined with natural variability of aggregate sources, leads to the need for identification of D-cracking-susceptible aggregates before they are used in concrete exposed to moisture and cycles of freezing and thawing.

D-cracking has been known since the 1930s (Stark and Klieger 1973); wide ranges of tests have been developed to try to identify susceptible aggregates. The most common proce-dure for identifying susceptible aggregates is ASTM C666/C666M. Concrete specimens made with the aggregate in ques-tion 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 beneficiation—A variety of techniques have been proposed to improve the performance of suscep-tible aggregates. Limiting the nominal maximum size was discussed in 4.3.2.1. Schwartz (1987) summarized other methods, including coating the aggregates to prevent their absorption of water, heavy media separation, and blending 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 of existing D-cracking—A considerable amount of concrete containing D-cracking-susceptible aggre-gates 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 significantly reduce the service life of 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 12 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 1994).

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 difficult 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 pave-ment that contains D-cracking-susceptible aggregates could be prevented by covering it with a sufficient thickness of asphaltic concrete. Janssen et al. (1986) 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 actu-ally accelerate the rate of D-cracking due to the increased potential for moisture migration during slower freezing rates. More than 6 in. (150 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. 1986; Janssen and Snyder 1994).

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 move-ment 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 pave-ment section in Ohio in 1992 (Janssen and Snyder 1994). Though initial laboratory testing indicated that sealer treat-ment delayed the resumption of D-cracking, field moni-toring showed that the D-cracking reappeared after six years (Janssen 2001).

4.3.4 Theories and mechanisms of D-cracking—Theo-ries 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, therefore, will concentrate on the characteristics of specific aggregates that make them susceptible, while other aggre-gates of the same type are not susceptible.

4.3.4.1 Pore size and size distribution—Kaneuji et al. (1980) 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 (1982) found that with one exception, the D-cracking-susceptible aggregates analyzed exhibited a predominance of pore sizes of 1.5 × 10–6 to 8 × 10–6 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 (1985) noted a reduced resistance to D-cracking for some aggregates when pavements containing susceptible aggregates were exposed to deicing salt. Other aggregates that they exam-ined showed no effect. A possible explanation is a change in pore structure due to etching of the pore walls by the salt.



Such action has been identified for aggregates containing calcite (Gillott 1978). This possible deicing salt effect should be noted when using field performance records to evaluate the D-cracking potential of an aggregate. Satisfactory perfor-mance in a pavement that never received deicing salt may not ensure the same performances where deicing salts are used.

CHAPTER 5—ALKALI-AGGREGATE REACTION

5.1—IntroductionThis chapter presents guidelines for minimizing the risk

of damaging expansion caused by alkali-aggregate reac-tion (AAR) in concrete construction. This risk may include potential damage from alkali-carbonate reactive or alkali-silica reactive aggregates. Procedures are discussed for eval-uating aggregates and selecting appropriate measures for controlling expansion are discussed. ACI 221.1R provides more detailed information on types of reaction, reaction mechanisms, reactive rock types, methods of testing aggre-gates, and preventive measures.

Alkali-silica reaction (ASR) was first detected in the late 1930s and first reported in 1940 (Stanton 1940a,b). Since its initial detection, ASR has been identified as a major cause of premature concrete deterioration throughout the world. Since Stanton’s first investigations, a wide range of testing 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 LiNO3. The reliability of these techniques varies and depends on, to some extent, the nature of the aggregate being tested and the testing environment.

The first case of alkali-carbonate reaction (ACR) was observed in Ontario, Canada, in the late 1950s (Swenson 1957). ACR occurs between alkali hydroxides and certain argillaceous dolomitic limestones. This reaction is charac-terized by rapid expansion and extensive cracking of the affected concrete. Structures suffering from ACR gener-ally 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 region-ally. Many specifications do not permit the use of reactive aggregates. When reactive aggregates are used in concrete, recommended preventive measures include limiting the alkali content of the concrete; using SCMs such as fly ash, slag cement, silica fume, or natural pozzolans; using lithium nitrate; or combining these methods.

A history of satisfactory field performance may be the most effective method for evaluating the potential for an

aggregate to cause AAR (5.3.1). Where such satisfactory field performance 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 absence of such field performance 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 difficult, or economically unfeasible, to control expansion with such materials.

5.2—Types of reactionsTwo types of AAR have been recognized: ACR and ASR.

Alkali-carbonate reaction is associated with the use of certain argillaceous dolomitic limestones. Confirmed cases of 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 reac-tion between an alkali source and certain calcium-magne-sium 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 field. Reactions involving such silica 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 10 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; these dolomites are characterized by a matrix of fine 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



Ca·Mg(CO3)2 + 2ROH → CaCO3 + Mg(OH)2 + R2CO3

(5.2.1a) (dolomite + alkali hydroxide → calcite + brucite + alkali carbonate)

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 1964; Tang et al. 1987; 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 dedolomi-tization (brucite and calcite)

The alkali carbonate produced in the dedolomitization reaction may react with calcium hydroxide (CH) in the cement paste as follows

R2CO3 + Ca(OH)2 → CaCO3 + 2ROH (5.2.1b)

thereby regenerating alkalis for further reaction. Thus, provided there is sufficient 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 ASR background—Alkali-silica reaction has been a known cause of concrete deterioration for more than 70 years. Since its discovery in the late 1930s (Stanton 1940a,

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 mech-anism by which the alkali-silica gel causes expansion and subsequent cracking in concrete is not yet entirely under-stood by researchers in the field.

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 main-tain 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: sufficient alkali, reactive silica, and adequate moisture.

5.2.2.1 Alkalis—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 Na2Oeq.

Equation (5.2.2.1) is used to determine the sodium oxide equivalency in the portland cement

Na2Oeq = Na2O + 0.658K2O (5.2.2.1)

where Na2Oeq is the total sodium oxide equivalent, in percent by mass; Na2O is sodium oxide content, in percent by mass; and K2O is potassium oxide content, in percent by mass.

The concentration of alkalis in portland cement generally ranges from 0.2 to 1.3 percent Na2Oeq, 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 Na2Oeq (Stanton 1940b). 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 effec-tively 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 (1983a) reports that the OH– ion threshold concentration for ASR is unlikely to be less than 0.25M and Kollek et al. (1986) suggest a threshold of

Table 5.2—Some examples of rock types and minerals susceptible to ASR

Reactive rocks Reactive minerals

Shale Opal

Sandstone Tridymite

Silicified carbonate rock Cristobalite

Chert Volcanic glass

Flint Cryptocrystalline (or microcystalline) quartz

Quartzite Strained quartz

Quartz-arenite

Gneiss

Argillite

Granite

Greywacke

Siltstone

Arenite

Arkose

Hornfels



0.20M; this equates to a pH threshold of 13.2 to 13.3 for ASR to occur.

5.2.2.2 Reactive silica—The degree of reactivity of aggre-gates depends on a number of factors, including the miner-alogy 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 suscep-tible 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 final 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 mois-ture to drive the expansion of the alkali-silica gel and sustain the reaction (Pedneault 1996).

5.2.2.4 Mechanism of 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 concentra-tion 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 1981) as follows

≡Si-OH + OH– → ≡Si-O- + H2O (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 1981)

≡Si-O-Si≡ + OH– → 2 ≡Si-O- + H2O (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 sufficient reserves of 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 of calcium—Bleszynski and Thomas (1998) and Thomas (2006a) concluded that significant expansion only occurs when an adequate supply of calcium is avail-able 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

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 1944)

b) Calcium hydroxide may act as a buffer maintaining a high level of OH– in solution (Wang and Gillott 1991)

c) High calcium concentrations in the pore solution prevent the diffusion of silica away from reacting aggregate particles (Bleszynski and Thomas 1998; Chatterji 1979; Chatterji and Clausson-Kaas 1984)

d) If calcium is not available, reactive silica may merely dissolve in alkali hydroxide solution without causing damage (Thomas 2006a; Diamond 1989)

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. 1991; Bleszynski and Thomas 1998)

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 of gel expansion—Although the mecha-nisms behind the formation of the gel are well understood, the actual mechanism for expansion of gel remains uncer-tain. 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 semi-permeable membrane, preventing the presence of large sili-cate 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 solu-tion through the cement paste, resulting in continued forma-tion 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 1944).

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 aggre-gate 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 SiO2 into the pore solu-tion. As the solvated hydroxyl and alkali ions infiltrate 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 SiO2 can diffuse out, while additional materials penetrate into the aggregate struc-ture. This imbalance results in an expansive force within the aggregate grain. If, however, the Ca(OH)2 and alkali



salt concentrations are low, the rate of penetration of the hydroxyl and alkali ions is also low, while the migration of SiO2 is increased, thus resulting in a lower expansive force (Chatterji et al. 1987).

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 fills more space 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 1995).

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 electri-fies the interface. This excess charge alters the properties of both the solid and liquid materials. Alkali-silica reaction involves the interaction of the highly-charged silica aggre-gate 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, nega-tively 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, potassium, and calcium ions, which surround the negatively charged silica surface. Once this double layer has been estab-lished, 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. 1997).

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 accu-rately explained the mechanism. For example, none of the four main theories discussed previously considers the poten-tially crucial role that short-range forces might play in expan-sion. Neither the osmotic pressure theory nor double-layer theory considers the potential effect of calcium on expan-sion. 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 alkali-aggregate reactivity

5.3.1 Field performance—A history of satisfactory field performance is possibly the best method for evaluating the potential for an aggregate to cause AAR. A number of factors have to be considered when analyzing field perfor-mance 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 field concrete as proposed in the new structure.

b) The concrete examined should be at least 10 years old and preferably more than 15 to 20 years old.

c) The exposure conditions of the field concrete should be at least as severe as those in the proposed structure.

d) In the absence of documentation conclusively demon-strating that the aggregate to be used in the proposed struc-ture is sufficiently similar to the field 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 used in the field structure should be considered.

f) Provided that satisfactory field 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 should be the first step 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 sufficient evidence to reject an aggregate on the basis of potential alkali reactivity or require that suitable preven-tive measures be explored and implemented. Generally, the examination cannot predict whether the type and distribu-tion of reactive minerals present will cause damaging expan-sion 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 result of incorrect material classification, but more likely is due to a failure to recognize certain minerals as potentially reactive (Rogers and Hooton 1991). Furthermore, the reac-tive constituents of some rocks may not be readily identified by optical microscopy.

5.3.3 Laboratory tests to identify 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 dura-tion 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



reactive of cement-aggregate combinations and is not recom-mended for evaluating ACR aggregates. Mortar bars (aggre-gate/cement = 2.25) are monitored for expansion under storage conditions at 100°F (38°C) in storage containers designed to maintain high humidity. For testing, in ASTM C227 coarse aggregates are crushed to a fine aggregate size fraction and both coarse and fine aggregates follow specific gradation requirements set forth in ASTM C227. The alkali content of the cement is not specified, 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 specification for concrete aggregates (ASTM C33/C33M) regards as reactive those cement-aggregate combinations that have an expansion greater than 0.05 percent at 3 months or 0.10 percent at 6 months. More stringent limits have been applied by certain agencies. For example, the U.S. Bureau of Reclamation requires specific combinations to expand by less than 0.05 percent at 6 months, or 0.10 percent at 12 months.

There are a number of limitations associated with this test. For example, it is difficult 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, espe-cially 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 compo-nent of the aggregate reacts relatively slowly. This has been observed in argillites and greywackes in which the reac-tive component is microcrystalline quartz (Grattan-Bellew 1989; Rogers and Hooton 1991). The consequences of alkali leaching may be reduced by storing bars over water without wicks and by raising the equivalent alkali content or Na2Oeq (note: Na2Oeq = Na2O + 0.658 K2O) of the cement to 1.0 percent, but expansions will still be less than 0.10 percent at 6 months with some reactive aggregates. Due to these limi-tations, 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. 100 (150 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 reac-tivity is poor for a number of reasons. Other mineral phases present in the aggregate may reduce the dissolved silica by precipitation (Bérubé and Fournier 1992a). 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 temper-ature 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 field performance being

classified as deleterious (Bérubé and Fournier 1993). Note that ASTM C289 was withdrawn by ASTM in 2016.

5.3.3.3 Accelerated mortar bar test (ASTM C1260)—This test is essentially the same as that developed by Oberhol-ster and Davies (1986). In this method, the length change of mortar bars (measuring nominally 1 x 1 x 11.25 in. [25 x 25 x 285 mm]) stored in 1 molar NaOH solution at 176°F (80°C) is monitored for 14 days. For testing, in ASTM C1260, coarse aggregates are crushed to a fine aggregate size fraction and both coarse and fine aggregates follow specific gradation requirements set forth in ASTM C1260. 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 (100°F [38°C] at 100 percent humidity) after 1 year (Oberholster and Davies 1986; Hooton and Rogers 1989). The test has been successfully used to identify alkali-silica reactive aggregates from across Canada (Grattan-Bellew 1989; Hooton and Rogers 1989, 1992; Bérubé and Fournier 1992b; Durand et al. 1990; Hooton 1991) and the United States (Stark et al. 1993), but is not considered suit-able for evaluating reactive alkali-carbonate rocks.

Interpretation of results is not simple, and various expan-sion criteria have been suggested (Thomas et al. 2007). Bérubé and Fournier (1992b) proposed a limit of 0.10 percent expansion after 14 days in 1M NaOH for quarried silicate and siliceous carbonate aggregates, and a limit of 0.20 percent for natural sands and gravels. Many aggre-gates with satisfactory field performance, however, produce expansions in excess of 0.25 percent in this test (Bérubé and Fournier 1992b). Consequently, aggregates should not be rejected on the basis of this test unless petrographic exami-nation confirms 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 0.10 percent at 14 or 28 days, yet cause deleterious expansion in ASTM C1293 and in the field (Ideker et al. 2012). In summary, this test should be used with caution, owing to the many poten-tial discrepancies between the performance of aggregates in this test and in the field (Thomas et al. 2007).

5.3.3.4 Concrete prism test (ASTM C1293)—This test is considered the most reliable for correctly identifying alkali-silica reactive aggregates (Thomas et al. 2006). A concrete mixture is proportioned with a cement content of 708 ± 17 lb/yd3 (420 ± 10 kg/m3) using a portland cement with an equivalent alkali content (Na2Oe) of 0.90 ± 0.10 percent. Sodium hydroxide is then added to the mixing water to provide a total alkali loading of the concrete of 8.85 lb/yd3 (5.25 kg/m3). This high alkali loading is necessary to induce expansion of slowly reactive rocks such as greywacke and argillites (Magni et al. 1987). Prisms are stored at 100°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 specified in both Canada and France (Thomas et al. 1997) and is recommended in ASTM C1293. This expansion corre-lates approximately to the point where cracking and signs of distress are first observed on the prisms. It also relates well



to field performance. 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 struc-tures, such as nuclear containment, or dimensionally sensi-tive structures, such as hydraulic dams, where small expan-sions may result in relatively large movements.

This test method has distinct advantages over the mortar bar tests (ASTM C227; ASTM C1260), as coarse aggregates can be tested without crushing to sand sizes. Furthermore, the larger test specimen reduces the effect of alkali leaching. The 12-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 undesir-able side effects (Ideker et al. 2012).

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 1990s to accelerate the test by increasing the temperature from the standard 100 to 140°F (38 to 60°C) (Ranc and Debray 1992; Murdock and Blanchette 1994; DeGrosbois and Fontaine 2000; Touma et al. 2001).

Unfortunately, simply raising the temperature and identi-fying expansion limits was not enough to produce a consis-tent and reliable test method. Fournier et al. (2004) iden-tified 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 selection of the nonreactive fine aggregate; they also proposed expansion limits. Further work by Ideker et al. (2010) made clear the profound effect of the selection of the nonreactive fine aggregate on expansions 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 C1293 (Ideker et al. 2010). Thus, the accelerated concrete prism test is not recommended for assessing aggregate reactivity or for deter-mining the efficacy of mitigation measures.

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. (1998, 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 160 mm) are cast, cured for 24 hours, and then soaked in water at 176°F (80°C) for an additional 24 hours. The bars are then moved into a 1 N NaOH at 176°F (80°C) and length change is moni-tored for 28 days. Aggregates are graded (or crushed where appropriate) to produce to a size fraction in the range of 0.2 to 0.4 in. (5 to 10 mm) and cast into mortar bars at a fixed 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 C1260). This accelerated test method is gaining momentum, and other researchers have shown its efficacy in testing poten-tially alkali-silica reactive coarse aggregates. This method may serve as a complementary test method, especially when ASTM C1260 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 identify 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 1M 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 0.10 percent at 1 month are generally taken to indi-cate 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 C1105) is recommended if the rock cylinder expansion exceeds 0.10 percent.

5.3.4.2 Chemical composition (CSA A23.2-26A)—The determination of potential ACR by chemical composi-tion involves analysis for CaO, MgO, and Al2O3 (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 the potentially expansive area of a CaO/MgO-versus-Al2O3 plot, and such aggregates should be tested by ASTM C1105. This test has helped to remove some of the difficulty in identifying reactive dolomitic limestones by petrographic examination.

5.3.4.3 Concrete prism test (ASTM C1105)—This test is similar to ASTM C1293 used to assess ASR (5.3.3.4), except for differences in storage temperature and the alkali content of the concrete. ASTM C1105 requires the testing to be carried out using a specific concrete mixture, with specimens stored at 73°F (23°C). Potentially deleterious reactivity is indicated if the expansion exceeds 0.015 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 infor-mation about the specific cement-aggregate combination tested, and that the absence of significant expansion in this test does not necessarily indicate that the aggregate is nonre-active. For instance, deleterious expansion may occur if the aggregate is used in concrete with a higher alkali content. CSA A23.2-14A requires potentially alkali-carbonate reac-tive rocks to be tested in concrete prisms under the same conditions as those used for ASR (that is, with 8.85 lb/yd3 [5.25 kg/m3] Na2Oe and stored at 100°F [38°C]). Further-more, the same expansion criteria are applied; aggregates are deemed to be reactive if the expansion exceeds 0.040 percent at 12 months, with no criteria for earlier test results. CSA A23.2-14A is aimed at establishing the reactivity of the rock and not assessing the performance of a particular cement-aggregate combination.



5.3.4.4 Chinese accelerated concrete microbar test (for alkali-carbonate reactive aggregates)—The Chinese accel-erated concrete microbar test was created to capture reac-tivity of alkali-carbonate rocks (Lu et al. 2008; Xu et al. 1998, 2000; Sommer et al. 2005). Further work by Lu et al. (2008) has shown the ability of this test to detect alkali-carbonate reactive aggregates across a range of aggregate gradations (0.2 to 0.4 in. [5.0 to 10.0 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 reli-able results than the accelerated concrete microbar test. This test is an attractive alternative to the longer-duration ASTM C1105 test for detecting ACR.

5.4—Preventive measuresThere are a number of preventive measures that can be used

to minimize the risk of damage due to ASR, which include:a) Using nonreactive aggregatesb) Limiting the alkali content of the concretec) Incorporating SCMsd) Using chemical admixtures, namely, lithium compoundsThese 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 1992).

5.4.1 Use of nonreactive aggregate—This approach is perhaps the most obvious and certain way of avoiding damaging reaction in concrete structures. Nonreac-tive 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 stringent to reflect the increasing number of aggregates implicated in field cases of 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 envi-ronments such as external source of alkalis like seawater or deicing salts, high cement contents, or extended service life.

5.4.2 Limiting alkali content of concrete—Stanton’s (1940a,b) work on AAR indicated that expansive reaction is unlikely to occur when the alkali content of the cement is below 0.60 percent Na2Oe. This value has become the accepted maximum limit for cement to be used with reac-tive aggregates in the United States, and appears in ASTM

C150/C150M 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 of potential reactivity. Some national specifications recognize this fact by specifying a maximum alkali content in the concrete; this limit was reported (Nixon and Sims 1992) to range from 4.21 to 7.58 lb/yd3 (2.5 to 4.5 kg/m3) Na2Oe. In some countries, the limit may vary depending on the reactivity of the aggregate (Oberholster 1994). In CSA A23.2-27A, the limit ranges between 3.03 to 5.05 lb/yd3 (1.2 and 3.0 kg/m3) Na2Oe. 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 sufficient safeguard in all cases. Stark (1980) reported damaging AAR in highway structures constructed using cements with alkalis in the range 0.45 to 0.57 percent Na2Oe. The reactivity of certain aggregates with low-alkali cements was confirmed in laboratory mortar bar expansion tests. Lane (1987) reported that some aggre-gates, classed as innocuous after 6 or 12 months in ASTM C227 with low-alkali cement (0.54 percent Na2Oe), showed delayed expansion and cracking after longer periods. Thomas (1996) reported evidence of ASR in a number of hydraulic structures with alkali contents below 4.0 lb/yd3 (2.4 kg/m3) Na2Oe.

Aggregates that are not normally reactive when used in concrete with low-alkali cement may be deleteriously reac-tive 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 (1978) reported increases in soluble alkali from 1.85 to 6.07 lb/yd3 (1.1 to 3.6 kg/m3) Na2Oe close to the surface of some highway structures. Migration of alkalis due to moisture, temperature, and electrical gradients has also been demon-strated in laboratory studies (Nixon et al. 1979; Xu and Hooton 1993).

There are many aggregates containing alkalis that may be leached out into the concrete pore solution, thereby increasing the risk of AAR (Stark 1980; Stark and Bhatty 1986; Way and Cole 1982; van Aardt and Visser 1977; Thomas et al. 1991; Bérubé et al. 2002). Stark and Bhatty (1986) reported that in extreme circumstances, some aggre-gates release alkalis equivalent to 10 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. (1987) 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 (1992) 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. 1987) have shown that exposure to NaCl solution and other alkali salts can cause considerable expan-sion and cracking in concrete.



5.4.3 Use of SCMs—In his second published article on ASR, Stanton (1940b) reported that expansion due to the reaction could be reduced by the use of pozzolanic cement containing finely ground shale or by replacement of 25 percent of high-alkali portland cement with pumicite. Subse-quent tests by Stanton (1950) confirmed the beneficial effect of a wide range of natural pozzolans and demonstrated that partially replacing portland cement with a sufficient 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 beneficial action of the pozzolan extended beyond merely diluting the cement alkalis. The first major use of a natural pozzolan to control ASR dates back to the 1940s when calcined siliceous shale, also called Puente shale, was used in the Davis Dam, together with low-alkali cement (Gilliland and Moran 1949). In the early 1950s, various studies (Cox et al. 1950; Barona de la O 1951; Buck et al. 1953) showed that other SCMs (namely, fly ash and slag) were also effective in reducing expansion. Later research showed that silica fume was highly efficacious in control-ling ASR with levels of 10 percent or less being sufficient to suppress damaging expansion in mortar bars with reac-tive aggregate from Iceland (Asgeirsson and Gudmundsson 1979) and South Africa (Oberholster and Westra 1981).

All SCMs contain alkalis and some, like fly ash, may contain substantially more than the portland cement they replace. This has led to considerable controversy in the past regarding whether the alkalis 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 reac-tion 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 of hydration, or fixed by the products of alkali-silica gel. In systems free of reac-tive aggregate, the partition of the alkalis between the pore solution and the hydrates is largely a function of the binder composition (Thomas 2011).

There is a large body of data in the literature that shows how various SCMs affect the composition of the pore solu-tion extracted from hydrated cement pastes (Longuet 1976; Diamond and López-Flores 1981; Diamond 1981, 1983a,b; Glasser and Marr 1985; Canham 1987; Canham et al. 1987; Kollek et al. 1986; Duchesne and Bérubé 1992, 1994; Kawamura and Takemoto 1988; Page and Vennesland 1983; Kawamura et al. 1987; Andersson et al. 1989; Durand et al. 1990; Yilmaz and Glasser 1990; Rasheeduzzafar and Hussain 1991; Shayan et al. 1993; Wiens et al. 1995; Nagataki and Wu 1995; Shehata et al. 1999; Ramlochan et al. 2000; Shehata and Thomas 2002; Bleszynski 2002; Boddy et al. 2003). Studies on the effect of fly ash and slag on the pore solution of pastes have been reviewed by Thomas (1996) and studies involving silica fume have been reviewed by Thomas and Bleszynski (2001). These studies show that the incorpora-tion of most SCMs leads to a reduction in the concentration

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. [1999], Ramlochan et al. [2000], Bleszynski [2002], and Shehata and Thomas [2002]). The most efficient SCM in terms of reducing the pore solution alkalinity is silica fume, followed closely by metkaolin and low-calcium fly ash. Slag and high-calcium fly ash are less effective and have to be used at higher cement replacement levels. Thomas (2011) showed that the concentration of alkali in the pore solution is a function of the composition of the binder (cement + SCM), especially its alkali (Na2Oe), 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 (Na2Oe·CaO)/(SiO2)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 efficacious in terms of controlling the alkali available for reaction.

Fig. 5.4.3a‒–Evolution of the pore solution in pastes containing SCM (Shehata et al. 1999; Ramlochan et al. 2000; Bleszynski 2002; Shehata and Thomas 2002).

Fig. 5.4.3b–‒Effect of 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).



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 1978; Rayment 1982; Uchikawa et al. 1989; Thomas et al. 1991). Glasser and Marr (1985) 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 posi-tive and the C-S-H tends to repel cations. As the Ca/Si ratio decreases, the positive charge decreases, becoming nega-tive at Ca/Si ratios less than 1.3 (Glasser 1992). Negatively charged C-S-H has an increased capacity to absorb cations, especially alkalis. Hong and Glasser (1999) confirmed 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 intro-ducing alumina into the C-S-H to form C-A-S-H (Hong and Glasser 2002).

SCMs that are highly efficient 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 C1293) 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 fly ash are most effective and have to be used at replacement rates of 10 to 30 percent, whereas slag and high-calcium fly 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 replace-ment. 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: Gener-ally 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 sufficient to control expansion due to ASR:

a) Silica fume: 10 to 15 percentb) Metakaolin: 15 to 20 percentc) Low-CaO fly ash: 20 to 30 percentd) Slag: 35 to 50 percente) High-CaO fly ash: greater than or equal to 40 percentAs 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 prescrip-tive guidelines developed from empirical data. AASHTO PP065 and ASTM C1778 provide both performance-based and prescription-based methodologies for determining the required SCM content.

5.4.4 Use of 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 fluoride, and alkyl alkoxy silane. Interest in the use of lithium compounds, specifically lithium nitrate, has resulted in significant research and testing in both laboratory and field 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 brief discussion of the findings regarding lithium salts follows.

5.4.4.1 Lithium salts—Although McCoy and Caldwell (1951) reported on the ability of lithium compounds (LiF, LiCl, and Li2CO3) 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 field trials, was conducted by Stark et al. (1993).

Initial work by McCoy and Caldwell (1951) and Lawrence and Vivian (1961) 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

Fig. 5.4.3c–‒Relationship between pore solution composi-tion and the chemical composition of the binder.

Fig. 5.4.3d–‒Effect of SCMs on 2-year expansion of concrete 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).



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 of 0.60 to 1.1 are sufficient to suppress expansion for many aggregate types (Sakaguchi et al. 1989; Stark et al. 1993; 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 sufficient. Additional caution is advised because insufficient lithium may actually increase expansion (Stark et al. 1993). Other research has shown that lithium salts such as LiOH and LiCO3 are less effective than LiNO3 in reducing or eliminating ASR (Collins et al. 2004). Several documents provide more detailed review material and guid-ance 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 1960), sodium silicofluoride, and alkyl alkoxy silane (Ohama et al. 1989). A wide range of compounds was studied by Hudec and Larbi (1989), but the results were largely inconclusive. Before such admixtures are recommended for commercial applica-tions, further research is required to confirm their efficacy and to elucidate their role in controlling ASR.

5.5—Tests for evaluating preventive measures5.5.1 ASTM C441/C441M—The pyrex mortar bar test, or

ASTM C441/C441M, has been commonly used for evalu-ating the efficacy of pozzolans and slag in controlling expan-sion due to ASR. This test method was developed in the 1940s as a method for assessing the suitability of pozzolans for use in concrete containing reactive aggregate, such as in connection with specifications for siliceous admixtures for the Davis Dam (Gilliland and Moran 1949). Early tests (Buck et al. 1953; Pepper 1964; Blanks 1950) indicated that fly ash and slag were less effective than highly siliceous natural pozzolans, and that they should be used in propor-tions exceeding 40 percent to be effective as defined by ASTM C441/C441M. Since then, numerous other workers have used this test to evaluate the performance of pozzolans and slag.

In ASTM C441/C441M, the 14-day expansion of mortar bars made with high-alkali cement (0.95 to 1.05 percent Na2Oe) and 25 percent fly ash by volume, or 50 percent slag and stored at 100°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 C618 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 of the field cement used will not exceed that of the test cement by more than 0.05 percent.

Although ASTM C989/C989M, a specification for slag, does not include a requirement relating to ASR, a nonman-datory appendix suggests the use of ASTM C441/C441M with 14-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 pozzo-lans or slag in this test have been criticized as too conserva-tive (Klieger and Gebler 1987; Kennerley 1988; Kennerley et al. 1981; Sturrup et al. 1983).

The potential for silica fume to reduce ASR expansion is apparent if used at the specified 25 percent by volume replacement in ASTM C441/C441M, with shrinkage often being observed after the normal 14-day testing period (Popovic et al. 1984). Perry and Gillott (1985) used various silica fume contents and found that 10 percent was effective in reducing the 14-day expansion by more than 75 percent compared with control (this was the acceptance criterion at the time). Other workers have confirmed the ability of 10 percent silica fume (Rasheeduzzafar and Hussain 1991; Hooton 1993) or less (Bérubé and Duchesne 1992) to meet this criterion. Perry and Gillott (1985), however, observed continued expansion of the silica fume specimens beyond 14 days and questioned the reliability of short-term testing by this method.

Results from tests with borosilicate glass (ASTM C441/C441M) have shown the effects of fly ash and slag to vary considerably between studies. The effect of fly ash has been shown to depend on its alkali content, calcium content, pozzolanicity, and fineness. The only consensus from the literature is that the effectiveness of fly ash and slag increases as the level of replacement increases, and that all fly ashes and slags can be used to control reaction, provided that they are used in sufficient quantity. Pepper and Mather (1959) reported the effectiveness of a pozzolan or slag was related to fineness, alkali release, and the amount of silica dissolved.

Borosilicate glass is extremely sensitive to test conditions (surface area, alkali content, and temperature) and contains significant quantities of alkalis that may be released into the pore solution. Furthermore, borosilicate glass produces damaging reaction in just a few days. Consequently, deter-mining 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/C441M (Hobbs 1989) because the results do not correlate well with data from concrete tests using natural aggregates (Bérubé and Duchesne 1992). Generally, the replacement level required to limit expansion in the boro-silicate glass mortar bar test is significantly higher than that required to limit expansion in concrete containing natural reactive aggregates; few commercial aggregates are as reac-tive as borosilicate glass.

ASTM C441/C441M is not recommended to determine the efficacy of 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



severely limits its ability to reliably predict the efficacy of lithium nitrate or other lithium compounds.

5.5.2 Accelerated mortar bar test (ASTM C1567)—ASTM C1567 was adopted in 2004 to assess the ability of SCMs to control ASR in mortar bars. The test method is essentially identical to ASTM C1260 (5.3.3.3) with the exception that a portion of the cement is replaced by the SCMs under test. The expansion limit of 0.10 percent at 14 days is specified in ASTM C1567, as it has been correlated with field perfor-mance of concrete (Thomas et al. 2006). While several agen-cies specify different expansion criteria in terms of percent expansion reached and test measurement age (14 versus 28 days), no additional information has yet been published to show a stronger correlation to the more reliable ASTM C1293 or field experience with similar combinations of aggregates and SCMs.

ASTM C1567 is not suitable for testing the efficacy of lithium nitrate, as a significant amount of the originally incorporated LiNO3 is leached to the aggressive 1 N NaOH host solution. Modified versions of ASTM C1567 for eval-uating lithium have been proposed (Tremblay et al. 2010; AASHTO PP065; USACE CRD-C 662).

5.5.3 Concrete prism test (ASTM C1293)—ASTM C1293 can be used to assess the efficacy of 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 C1293 as outlined in 5.3.4.3, with the exception that the duration is extended to 2 years. Measurements are typically taken at 3- or 6-month intervals after the first year of testing following the standard recommendation for concrete prism testing.

An expansion limit of 0.04 percent at 2 years is specified by CSA A23.2-27A and is recommended in the appendix to ASTM C1293. This expansion limit has shown a strong correlation to field structures cast with similar preventive measures. This method can also be used to effectively deter-mine the required dosage of lithium salts—namely, lithium nitrate—to mitigate deleterious ASR.

5.6—Protocols for minimizing the risk of alkali-aggregate 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 poten-tially 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 avail-able for selecting preventive measures. In 2010, the Cana-dian protocol was modified and adopted by AASHTO as AASHTO PP065. In 2014, AASHTO PP065 was modified and adopted as ASTM C1778.

The development of AASHTO PP065 is described by Thomas et al. (2008a). A flowchart describing the protocol for determining aggregate reactivity is presented in Fig. 5.6. Although aggregates can be accepted solely on the basis of satisfactory field 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 iden-tify a reactive aggregate. Laboratory expansion testing is recommended and the preferred test is ASTM C1293, 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 C1260, recognizing that it frequently results in false positives in that it identifies some nonreactive aggregates as reactive, and occasionally results in false negatives in that it fails to correctly identify some reactive aggregates. Quarried carbon-ates are evaluated on the basis of their chemical composition (MgO, CaO, and Al2O3) 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 C1293 or ASTM C1567. The dosage of lithium can also be determined using ASTM C1293 or a modified version of ASTM C1567. ASTM C1293 is preferred, but due to its 2-year duration, use of ASTM C1567 is also acceptable.

The practice also contains a protocol for determining the appropriate alkali limit or the level of SCM using a prescrip-tive 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 SCMThe maximum alkali content of the concrete varies within

the range from 3 to 5 lb/yd3 (1.8 to 3.0 kg/m3) Na2Oe, depending on the risk of ASR and the level of prevention required. Similarly, the minimum SCM content ranges from 15 percent fly ash or 25 percent slag where the risk of ASR is low and only mild mitigation measures are required, to 35 percent fly 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 100-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.



CHAPTER 6—SULFATE ATTACKSulfate 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 sulfate attack6.1.1 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 of sulfide or sulfate-bearing minerals. Indus-trial and agricultural effluents, as well as municipal waste-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. Soil fills containing

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 appro-priate preventive measures

Fig. 5.6‒–Sequence of test to determine aggregate reactivity (CSA A23.2-27A).



industrial waste products, such as slag from iron processing, can leach sulfate ions.

There are other environments where multiple deterio-ration 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 phys-ical 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 1968; Harboe 1982; Reading 1975, 1982; U.S. Bureau of Reclamation 1975; Verbeck 1968); the prairie provinces of Canada (Hamilton and Handegord 1968; Hurst 1968; Price and Peterson 1968); and the Middle East (French and Poole 1976). Other non-arid countries, such as England (Bessey and Lea 1953) and Norway (Bastiensen et al. 1957), have also experienced sulfate attack on concrete.

6.1.2 Historical background—Sulfate attack has been recognized for nearly 250 years. Smeaton (1791), requiring a material that would harden under water, developed the first hydraulic lime formulation of the industrial age, the precursor of modern portland cements. While developing a cement formulation for use in the construction of the Eddys-tone Lighthouse, Smeaton described its attack by sulfate-containing solutions in 1756. During the nineteenth century, various aspects of sulfate attack of concrete were studied (Bogue 1955).

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 (C3A) to form various sulfate-containing phases that adversely affect concrete dura-bility. Bates et al. (1913) 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 first described by Candlot in 1880, the precise compositions of calcium sulfoaluminate hydrates were not established until 1929. Lerch et al. (1929) established the composition of ettringite, known as Cand-lot’s salt, as 3CaO·Al2O3·3CaSO4·31H2O and that of mono-sulfate as 3CaO·Al2O3·CaSO4·12H2O. In addition to estab-lishing the compositions of these salts, Lerch et al. (1929) also showed that monosulfate converts to ettringite when an external source of sulfate is provided. This appears to be

the first 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. (1929), the role of C3A in sulfate attack had been recognized earlier. Concern about the C3A content of cements had stimulated a variety of studies of cement compositions over a period of 20 years or more. Bates and Klein (1917), who studied the proper-ties of calcium silicates and C3A, reached the conclusion that it would be impossible to commercially produce a portland cement containing less than approximately 1 percent alumina because of the high firing temperatures required. Work in Canada by Thorvaldsen, however, identified the means to reduce C3A contents by changing the proportions of C3A and C4AF 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 1933). Today, low-C3A cements are routinely produced. Sulfate attack, however, involves phenomena in addition to the formation of ettringite. Consequently, even so-called zero-C3A cements might not be immune to sulfate attack.

The basis of the most commonly used method for estab-lishing the sulfate resistance of cements by measurement of expansion should also be credited to Thorvaldson et al. (1927, 1929). 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 C1012/C1012M.

Resistance to sulfate attack is increased by controlling both cement composition and concrete permeability. The importance of this was demonstrated in studies by Verbeck (1968) and Stark (1989b) that showed that reduction of permeability was of greater importance in limiting sulfate attack than was using a sulfate-resistant cement composi-tion. In a 40-year summary of U.S. Bureau of Reclama-tion 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 C3A 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 significant damage can occur in the w/cm range from 0.45 to 0.53.

An appreciation that limitations on C3A content and on w/cm are both needed to produce sulfate-resistant concrete have been embodied in various codes and stan-dards governing the selection of concrete for use in sulfate environments. The U.S. Bureau of Reclamation (1975) has formally recognized these requirements since 1949.

Table 6.1.1––Mineral names and general composition often used in reports of sulfate attack

Anhydrite CaSO4 Thenardite Na2SO4

Bassanite CaSO4 ≈ 0.5 H2O Mirabilite Na2SO4 ≈ 10H2O

Gypsum CaSO4 ≈ 2H2O Arcanite K2SO4

Kieserite MgSO4 ≈ H2O Glauberite Na2Ca(SO4)2

Epsomite MgSO4 ≈ 7H2O Langbeinite K2Mg2(SO4)3



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 recog-nized in Brown (2002):

a) Ettringite (AFt), monosulfate (AFm), and gypsum formation

b) Sulfate-containing salt formation at or near an evapora-tive surface

c) Thaumasite formationMechanisms for these three forms of external sulfate attack

are described as follows, and publications discussing them in detail include Lea (1971), Mehta (1976, 1992), Mehta and Monteiro (2006), DePuy (1994), Taylor (1997), Hewlett (1998), and Skalny et al. (1998). Publications with particular emphasis on permeability and the ability of concrete to resist ingress and movement of water include Reinhardt (1997), Hearn et al. (2006), and Diamond (1998).

6.1.3.1 Sulfate attack associated with ettringite and gypsum formation—These forms of sulfate attack occur when external sulfate-containing solutions infiltrate 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 influenced by pH because ettringite is not stable below a pH of approximately 10 and gypsum is not stable above a pH of approximately 10.5 (Taylor 1997). This explains the common observations in both laboratory studies and in field concrete of gypsum near exposed concrete surfaces, where carbonation has lowered the pH, and of ettringite in the inte-rior of gypsum deposits.

In mature concrete, ettringite typically forms directly from the monosulfate that had formed during the hydra-tion 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 expan-sion; this damage mechanism is described in early studies of sulfate attack. However, ettringite may also form as secondary deposits in voids and cracks in mature field concrete exposed to wetting and drying. Consequently, the presence of ettringite in concrete does not necessarily indi-cate 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 mecha-nism involves chemical alterations that destabilize the C-S-H and calcium hydroxide (CH), and can result in the formation of microcracks without significant expansion.

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, CaSO4·2H2O) is generally the least aggressive because its solubility is significantly lower than that of sodium and magnesium sulfate. Calcium sulfate solutions can, however, attack concrete (Thorvaldson 1954; Taylor 1997; 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 1947), thus allowing the devel-opment 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 forma-tion of gypsum and ettringite within the cement paste at the expense of the normal cement hydration products of CH, monosulfate, hydrated C3As, 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 evap-orative 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 evapora-tive surfaces may indicate the occurrence of external chem-ical 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 partici-pate 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 1995; Taylor 1997). 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 solu-tion 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 mech-anism 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



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 interac-tion between sulfate ions and the hydrated phases of port-land cement. Physical salt attack can occur together with chemical sulfate attack.

If sulfate ions are identified in the soil, groundwater, or as part of the efflorescence on concrete surfaces, any damage of concrete by physical salt attack, including damage to archi-tectural surfaces, requires serious consideration and evalua-tion. Chapter 8 discusses physical salt attack.

6.1.3.3 Thaumasite formation—Although it has a different chemical composition, thaumasite has a similar crystal struc-ture to ettringite. While it is typical to denote ettringite by its oxide composition as 3CaO·Al2O3·3CaSO4·32H2O, the appropriate representation of the molecular structure of ettrin-gite is (Ca3Al(OH)6.12H2O3+)2(SO4

2–)3·2H2O·Ca3Al(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 substi-tuted for calcium; trivalent cations, such as transition metal ions, can substitute for aluminum; tetravalent ions, such as Si4+, can also substitute for Al3+. One such silica-based compound is thaumasite (CaO.SiO2·CaSO4·CaCO3·15H2O or CaSiO3·CaCO3·CaSO4·15H2O).

Thaumasite was first identified as occurring in deterio-rating concrete in 1965 by Erlin and Stark (1966), and later by Bickley et al. (1994), and has been extensively investi-gated by Matthews (1994), the Thaumasite Expert Group (1999) 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 (1966), however, found lath-shaped crystals where there were petal overgrowths of thaumasite on ettringite, finely 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 exclu-sively associated with its formation. Similarly, if a source of readily soluble CaCO3 is present within the concrete, thau-masite 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, alka-line conditions typically experienced by buried concrete

structures. The occurrence of thaumasite in deteriorated building materials has been identified in a number of 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 encoun-tered so far was in the Canadian Arctic (Bickley et al. 1994).

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 trans-forms the cement paste matrix into a white, soft, noncohe-sive mass. From the formula for thaumasite, it is seen that carbonate ions are also necessary.

A positive identification of 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 1999) 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 significant damage to the cement paste matrix of the concrete or mortar. The main hall-mark 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 filled 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 harm-lessly 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 1986; Regourd and Hornain 1986). 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 aggre-gates 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 (15°C). Although it can form at temperatures up to 77°F (25°C), the rate is much slower (Thaumasite Expert Group 1999; Alksnis and Alksne 1986).

6.1.4 Recommendations6.1.4.1 Sulfate attack associated with ettringite, mono-

sulfate, and gypsum formation—Protection against the various forms of sulfate attack is obtained by proportioning



concrete mixtures to minimize the ingress and move-ment of water using appropriate ingredients. The sulfate resistance of portland cement generally decreases with an increase in its calculated tricalcium-aluminate (C3A) content (Mather 1968; Stark 2002). Accordingly, ASTM C150/C150M includes Type V sulfate-resisting cement for which a maximum of 5 percent calculated C3A is permitted, and Type II moderately sulfate-resisting cement for which the calculated C3A 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. 1986). Therefore, ASTM C150/C150M provides that in Type V cement, the C4AF + 2C3A 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-expan-sion test (ASTM C452/C452M) can be used in place of the chemical requirements (Mather 1978). 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 C1012/C1012M is discussed by Patzias (1991). ASTM C1157/C1157M, ASTM C595/C595M, and CAN/CSA A3000 blended cements also use ASTM C1012/C1012M expansion limits to qualify moderate (MS) or high (HS) resistant performance. Both ACI 318 and CSA A23.1-14/CSA A23.2 allow the use of other combinations of cementing materials in sulfate exposure, provided that performance testing using ASTM C1012/C1012M demon-strates that the expansion limit for the appropriate exposure class (Table 6.1.4.1a) is not exceeded. Note that in Table 6.1.4.1b, performance testing is not required if ASTM C150/C150M 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 fluid ingress). Care should be taken to ensure that the concrete is designed and constructed to minimize shrinkage cracking. Proper place-ment, compaction, finishing, and curing of concrete are essential to minimize the ingress and movement of fluids that carry aggressive salts. Recommended procedures for these are found in ACI 304R, ACI 302.1R, 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 6.1.4.1b for exposures defined in Table 6.1.4.1a. Both recommenda-tions are important, as limiting only the type of cementi-tious material is not adequate for satisfactory resistance to sulfate attack (Kalousek et al. 1976; Stark 2002; Monteiro and Kurtis 2003).

Table 6.1.4.1b 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 effluents carried in concrete pipes.

The field conditions of concrete exposed to sulfate are numerous and variable. The aggressiveness of the condi-tions depends on, among other things, soil saturation, water movement, ambient temperature and humidity, concentra-tion of sulfate, and type of sulfate or combination of sulfates involved. Table 6.1.4.1b provides criteria that should maxi-mize 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 6.1.4.1a and 6.1.4.1b).

6.1.4.3 Thaumasite formation—Use of Types II or V sulfate-resisting cements does not prevent thaumasite sulfate 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 fines in aggregates or cements that are well above the concentra-tions allowed by ASTM C150/C150M; 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 specifi-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 C1012/C1012M, when modified 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. 2010); this test was adopted in CSA A3000 in 2010. 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 confirmed by Bellmann and Stark (2008) where mortar bars made with a European CEM IIIBi cement with 65 percent slag cement was resistant to thaumasite after exposure to sulfate solutions at 46°F (8°C). They attributed the good

Table 6.1.4.1a—Severity of exposure conditions determined from sulfates in soil or water

Exposure classWater-soluble sulfate

(SO42–)* in soil, %

Sulfate (SO42–)* in

water, ppm

S0 (not applicable) SO42– < 0.10 SO4

2– < 150

S1 (moderate) 0.10 ≤ SO42– < 0.20 150 ≤ SO4

2– < 1500, or seawater

S2 (severe) 0.20 ≤ SO42– ≤ 2.00 1500 ≤ SO4

2– ≤ 10,000

S3 (very severe) SO42– > 2.00 SO4

2– > 10,000*Sulfate expressed as SO4 is related to sulfate expressed as SO3, as given in reports of chemical analysis of portland cements as follows: SO3 × 1.2 = SO4.



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. 2010). Bellmann and Stark (2007) also found that when a CEMIIA-L cement, which is portland cement with up to 20 percent interground limestone, was replaced with either 20 or 40 percent Class F fly ash, there was no damage after 4.5 years of storage in 1500 mg/L SO4

= (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, representa-tive samples should be obtained of both the fluid and sulfate 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 C1580 is recom-mended (Hayes 2007). Although other methods have been used, the results are often affected by the test method, espe-cially 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 6.1.4.1b is permitted for any class of exposure. Any binary or ternary blend of portland cement of any type meeting ASTM C150/

C150M, ASTM C595/C595M, or ASTM C1157/C1157M with fly ash or natural pozzolan meeting ASTM C618, silica fume meeting ASTM C1240, or slag cement meeting ASTM C989/C989M is permitted if it meets the expansion limits in Table 6.1.4.1b when tested in accordance with ASTM C1012/C1012M.

The portland-cement portion of the test mixture should always consist of cement with Bogue-calculated C3A content of not less than that being proposed for use. Material quali-fication tests using the expansion limits in Table 6.1.4.1b 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 uniformity of pozzolans and slag cement—The proportion of fly 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 recom-mendations of Table 6.1.4.1b and Section 6.1.6. In blends or mixtures with portland cement containing only one SCM, such as fly ash, natural pozzolan, silica fume, or slag cement, the proportion of fly ash or natural pozzolan can generally be expected to range between 15 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 12 percent by mass of the total cementitious material, and the proportion of cement

Table 6.1.4.1b—Requirements to protect against damage to concrete by sulfate attack from external sources of sulfate

Severity of potentialexposure

w/cm by mass, maximum

Prescriptive cementitious material requirements Performance cementitious material requirements

Cement types*Maximum expansion when tested using ASTM

C1012/C1012M

ASTMC150/C150M

ASTMC595/C595M

ASTMC1157/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 0.10% — —

S2 0.45† Type V#IP (HS), IS (<70) (HS), IT (P<S<70) (HS), or IT

(P≥S) (HS)HS 0.05% 0.10%|| —

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†† — — 0.10%

*Alternative combinations of cementitious materials to those listed in Table 6.1.4.1b can be permitted when tested for sulfate resistance and meeting the ASTM C1012/C1012M expansion criteria for the severity of potential exposure.‡Other available types of cement, such as ASTM C150/C150M Type I or Type III can be permitted in Exposure Classes S1 if the C3A content is less than 8 percent.§For seawater exposure, other ASTM C150/C150M cement types with C3A contents up to 10 percent are permitted if w/cm does not exceed 0.40. (Refer to Section 6.3 on seawater exposure.)#An ASTM C150/C150M Type III cement with the optional limit of 5 percent can be permitted or ASTM C150/C150M cement of any type having expansion at 14 days no greater than 0.040 percent when tested by ASTM C452/C452M.||The 12-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 specified 28-day compressive strengths of 26, 29, and 33 MPa (3750, 4250, and 4750 psi), respectively.**As stated in ACI 318, the amount of the specific source of the pozzolan or slag cement to be used shall be at least the amount that has been determined by service record to improve sulfate resistance when used in concrete containing Type V cement. Alternatively, the amount of the specific source of the pozzolan or slag cement to be used shall be at least the amount tested in accordance with ASTM C1012 and meeting the criteria shown in the table.††For Exposure Class S3, ASTM C1157/C1157M HS cement must contain pozzolan cement, slag cement, or both.



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 uniformity of the fly ash or slag cement used in the project should be within the following limits compared to that used in the mixtures tested to meet the recommenda-tions of Table 6.1.4.1b and Section 6.1.6:

a) Fly ash: reported calcium-oxide content (analyzed in accordance with ASTM C114) no more than 2.0 percentage points higher than that of the fly ash used in the test mixture

b) Slag cement: reported aluminum-oxide content (analyzed in accordance with ASTM C114) 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 C3A content no higher than that used in the mixtures tested to meet the recommendations of Table 6.1.4.1b 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 significantly reduce the permeability of concrete (Bakker 1980; Mehta 1981). They also combine with the alkalis and CH released during the hydration of the cement (Vanden Bosch 1980; Roy and Idorn 1982; Idorn and Roy 1986), reducing the potential for gypsum formation (Lea 1971; Biczok 1972; Kalousek et al. 1972; Mehta 1976).

Table 6.1.4.1b 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 18-month ASTM C1012/C1012M expansion limit in Table 6.1.4.1b is met. Some pozzolans, especially Class C fly ashes, decrease the sulfate resistance of mortars in which they are used (Mather 1981b; Mather 1982). Good results were obtained when the pozzolan was a fly ash meeting the requirements of ASTM C618 Class F (Dikeou 1975; Dunstan 1976). Slag cement should meet ASTM C989/C989M and silica fume should meet ASTM C1240.

In concrete that is made with non-sulfate-resisting cements, calcium chloride reduces resistance to attack by sulfate (U.S. Bureau of Reclamation 1975) 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 accept-able amounts as an accelerating admixture (Mather 1992). Calcium chloride, however, can induce and accelerate corro-sion of reinforcing steel and aluminum conduit.

6.2—Internal sulfate attack6.2.1 Concrete materials—Cementitious materials

meeting current ASTM specifications will not have sulfate contents that are deleterious to concrete. Allowable sulfate contents in cements meeting ASTM C150/C150M were increased several times from 1941 to 1971 (Hooton 2008), as

cement compositions and finenesses changed to allow better optimization of sulfate contents. ASTM C150/C150M now allows SO3 limits to be exceeded if it can be demonstrated (typically using ASTM C563) that the optimum SO3 content is above the stated limit. In this case, ASTM C1038/C1038M must show that the SO3 content of the cement will not result in adverse expansions. These results are considered satisfac-tory when a 14-day expansion limit of 0.020 percent is spec-ified. This test and expansion limit has also been adopted in ASTM C1157/C1157M and is used for Canadian portland cements, blended cements, and combinations of cementi-tious 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 sulfides such as pyrite. Limits are placed on sulfides in air-cooled, blast-furnace slag aggregate. The sulfides in slag aggregate include iron sulfide and calcium sulfide. 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 C1602/C1602M. Sulfates in chemical admixtures meeting ASTM C260/C260M, ASTM C494/C494M, and ASTM C1017/C1017M will not be dele-terious to concrete.

6.2.2 Delayed ettringite formation (DEF)6.2.2.1 Occurrence—Under certain conditions, heat-

cured concrete elements can suffer expansion and cracking on subsequent exposure to moisture. This form of deterio-ration 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 1990s, 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 1987; Tepponen and Eriksson 1987; Vitousova 1991; Mielenz et al. 1995; Oberholster et al. 1992; Shayan and Quick 1992). In most, if not all, of the alleged cases of DEF, other mechanisms of deterioration, especially alkali-silica reaction (ASR) and freezing-and-thawing damage, have also been implicated, making it difficult to identify 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



been successful in eliminating damage due to DEF (Skalny and Locher 1997; Stark and Bollmann 1999).

Note that the risk of damage due to DEF is not restricted to heat-cured precast concrete elements. Internal concrete temperatures may increase sufficiently, 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 mois-ture 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 tempera-tures. This delayed formation of ettringite may, under some circumstances, result in expansion of the paste and conse-quent 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 100 nm) of the cement paste (Taylor et al. 2001; Lawrence 1995). The expansion of cement paste results in the formation of gaps around aggregate particles and cracking of the cement paste (Johansen et al. 1993). Ettringite eventually reprecipitates into these gaps and the cracks, but this is not a cause of damage (Johansen et al. 1993).

Reviews of laboratory studies (Day 1992; Thomas 2001) and recent data (Thomas et al. 2008b) indicate that expan-sion due to DEF is unlikely to occur unless mortar or concrete specimens are subjected to elevated temperatures in excess of approximately 160°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-fineness cements produced from clinker with high concentrations of C3A, C3S, and Na2Oeq, and consequently having a high SO3 content, have the greatest susceptibility to DEF expansion when heat cured. Kelham (1996) showed a clear correlation between the 2-day compressive strength of mortar and the expansion of the same mortar subsequent to heat curing at 194°F (90°C). Although studies have shown relationships between DEF expansion and various compositional param-eters of the cement, there is no single parameter that can

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 C3A, C3S, Na2Oeq, and SO3 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 1996; 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. 1980; 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 fly ash or slag cement.

6.2.2.3 Recommendations—To minimize the risk of poor durability due to deleterious DEF reactions associ-ated with exposure to elevated temperatures at early ages, the maximum internal temperature of concrete should be controlled such that it does not exceed 158°F (70oC) at any time. If temperatures in the range of 158°F < T ≤ 185°F (70 to 85°C) are unavoidable, the measures in Table 6.2.2.2 should be adopted.

6.3—Seawater and brine exposure6.3.1 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 flowing in from rivers (Hewlett 1998).

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 accompa-nying 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 1998).

These combined deleterious effects can cause severe defects in concrete in the course of a very few years, espe-cially in tropical climates where high temperature increases the rate of deterioration. This section focuses on the influ-



ence of sulfates in seawater. Section 7.2 describes the influ-ence 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 1966). Concrete in seawater often exhibits erosion, softening, or loss in mass as a result of sulfate attack as opposed to the expansion, which may also occur in nonsa-line 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 C3A 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 C3A in the cement to form chloro-aluminates, such as Friedel’s salt (Verbeck 1975). Formation of chloroaluminates does not result in undesirable expan-sion, and it also lowers the amount of C3A 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 reac-tions 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 1998). Sulfate attack in seawater can also lead to decalcification of C-S-H. The formation of 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 1998; 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 expo-sure are provided in Table 6.1.4.1b under S1 exposure.

The performance of concretes continuously immersed in seawater made with ASTM C150/C150M cements having C3A contents as high as 10 percent has proven satisfactory, provided the permeability of the concrete is low (Browne 1980). The U.S. Army Corps of Engineers (USACE) (1984) and the Portland Cement Association recommend up to 10 percent Bogue-calculated C3A for concrete that will be permanently submerged in seawater if the w/cm is kept below 0.45 by mass.

CSA A23.1-14/CSA A23.2 allows w/cm of up to 0.50 and recommends 8 percent maximum C3A content. With rein-forced 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 C3A content decreases resistance to chloride penetration; therefore, CSA A23.1-14/CSA A23.2 suggests that C3A contents of portland cement be kept between 4 and 8 percent to help protect the reinforcement by increasing chloride binding.

Verbeck (1968) and Regourd et al. (1980) showed, however, that there may be a considerable difference between the calculated and the measured phase composition of cement, especially as far as C3A and C4AF are concerned. Therefore, the interrelation between the measured C3A 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 357.1R for expo-

Table 6.2.2.2—Recommended measures for reducing potential for DEF in concrete exposed to elevated temperatures at early ages*

Maximum concrete temperature T Prevention required

T ≤ 158°F (70°C) No prevention required

158°F < T ≤ 185°F(70°C < T ≤ 85°C)

Use one of the following approaches to minimize the risk of expansion:1. Portland cement meeting requirements of ASTM C150/C150M moderate or high sulfate-resisting and low-alkali cement with a fineness value less than or equal to 430 m2/kg2. Portland cement with a 1-day mortar strength (ASTM C109/C109M) less than or equal to 2850 psi (20 MPa)3. Any ASTM C150/C150M portland cement in combination with the following proportions of pozzolan or slag cement: a) Greater than or equal to 25 percent fly ash meeting the requirements of ASTM C618 for Class F fly ash b) Greater than or equal to 35 percent fly ash meeting the requirements of ASTM C618 for Class C fly 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 C1240) in combination with at least 25 percent slag cement e) Greater than or equal to 5 percent silica fume (meeting ASTM C1240) in combination with at least 20 percent Class F fly ash f) Greater than or equal to 10 percent metakaolin meeting ASTM C6184. An ASTM C595/C595M or ASTM C1157/C1157M blended hydraulic cement with the same pozzolan or slag cement content as listed in Item 3

T > 185°F (85°C) The internal concrete temperature should not exceed 185°F (85°C) under any circumstances.*Assembled from Ghorab et al. (1980), Ramlochan et al. (2003), Thomas (2001), Thomas et al. (2008b).



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 357.1R 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 1/10 or 1/100 that of comparable concrete of equal strength made only with portland cement (Bakker 1980). The satis-factory performance of concretes containing slag cement in a marine environment has been described (Lea 1971; Vanden Bosch 1980; Mather 1981a).

Concrete should be designed and constructed to minimize crack widths, therefore limiting chloride penetration to rein-forcement and avoiding the concentration of sulfates. Addi-tionally, 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—GeneralConcrete is rarely attacked by chemicals in their solid

form. To produce a significant attack on concrete, aggressive 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 condi-tions where aggressive chemicals are present, some kinds of chemical environments will significantly shorten the service life of even the best concrete unless specific measures are taken.

An understanding of these exposure conditions permits measures to be taken to prevent or slow deterioration. Lower permeability will hinder the infiltration of aggressive chemi-cals. Less-reactive paste can be effective in mitigating dete-rioration. Concrete members exposed to aggressive solutions that are under hydraulic pressure from one side may be more

vulnerable because the hydraulic gradient can accelerate the infiltration of the aggressive solution into the concrete.

This chapter discusses aggressive chemical exposures, including: seawater, acids, fresh water, carbonation, indus-trial chemicals, deicing chemicals, and environmental structures. Some useful summaries of the potential effects of chemical exposures include Lea (1971), Biczok (1967), Scrivener and Young (1997), Hewlett (1998), Eglinton (1998), Portland Cement Association (2001), ACI 515.2R, and ACI 350.1. Table 7.1a summarizes the effects of the more common chemicals that lead to the deterioration of concrete. Table 7.1b summarizes factors that may affect the rate of chemical attack.

7.2—Seawater7.2.1 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, contin-uous 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 section mainly focuses on the influence of chloride and magnesium ions in seawater. Section 6.3 describes the influ-ence of sulfates in seawater.

7.2.2 Reaction mechanisms—The chemical reactions that affect concrete exposed to seawater can be complex. The most significant aggressive species include magne-sium, sulfate, and chloride. Magnesium ions may react with calcium hydroxide (CH) and form magnesium hydroxide or brucite (Lea 1971). This phase is highly insoluble and

Table 7.1a—Effect of common chemicals on concrete*

Rate of attack at ambient temperature Inorganic acids Organic acids Alkaline solutions Salt solutions Miscellaneous

Rapid Hydrochloric NitricSulfuric

AceticFormicLactic

— Aluminum chloride —

Moderate Phosphoric Tannic Sodium† hydroxide greater than 20 percent

Ammonium nitrateAmmonium sulfate

Sodium sulfateMagnesium sulfate

Calcium sulfate

Bromine (gas)Sulfite liquor

Slow Carbonic —Sodium† hydroxide

10 to 20 percent sodium hypochlorite

Ammonium chloride Magnesium chloride

Sodium cyanide

Chlorine (gas)SeawaterSoft water

Negligible — OxalicTartaric

Sodium hydroxide Less than 10 percent sodium hypochlorite

ammonium hydroxide

Calcium chlorideSodium chloride

Zinc nitrateSodium chromate

Ammonia (liquid)

*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.



may fill pores on the outer surface of the concrete, which can reduce permeability and limit further infiltration of aggressive waters. However, brucite is not always observed in seawater exposures. Magnesium can react with any hydrate normally 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 forma-tion 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 poten-tial for corrosion of embedded steel.

Mather (1966) 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 C3A content than would be expected of the same cement in a nonmarine exposure where the water has the same sulfate ion content but lacks significant chlo-ride concentration. Concretes made with portland cements having C3A contents as high as 10 percent and subject to continuous immersion in seawater have performed satis-factorily, provided that the permeability of the concrete is low (Browne 1980). USACE EM 1110-2-2000:1994 and the Portland Cement Association recommend up to 10 percent calculated C3A for concrete that will be permanently submerged in seawater if the w/cm is kept below 0.45 by mass. Verbeck (1968) and Regourd et al. (1980) showed that there may be a considerable difference between the calcu-lated and measured clinker composition of cement, particu-

larly with respect to the proportions of C3A and C3AF. There-fore, the interrelation between the measured C3A content and the resistance to seawater may be uncertain. CSA A23.1-14/CSA A23.2 recommends the use of cements with between 4 and 8 percent C3A unless slag or pozzolans are used.

7.2.3 Mitigation—The mitigation of seawater expo-sure 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 357.1R 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 consoli-dation, and adequate curing are important to achieve satis-factory performance.

The use of slag cement and silica fume may reduce the permeability and increase the performance of concrete in a marine environment (Lea 1971; Fidjestøl and Frearson 1994; Mather 1981a; Vanden Bosch 1980). 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 1980). Concretes made with combina-tions of cement and silica fume, and with combinations of cement, slag cement, and silica fume also have lower perme-ability and good performance in seawater exposure.

Concrete should be designed and constructed to minimize the length, width, and number of cracks to limit seawater

Table 7.1b—Factors influencing chemical attack on concreteFactors that accelerate or aggravate attack Factors that mitigate or delay attack

1. High permeability due to:a) High water absorptionb) High w/cmc) Poor consolidationd) Poor curinge) Cracking and microcracking

1. Low-permeability concrete* achieved by:a) Proper mixture proportioning†

b) Reduced unit water contentc) Increased cementitious material contentd) Appropriate use of supplementary cementitious materials (SCMs)e) Air entrainmentf) Adequate consolidationg) Effective curing‡

2. Cracks and separations due to:a) Loading/stress concentrationsb) Thermal stressc) Shrinkage

2. Reduced tensile stress in concrete by:#

a) Using tensile reinforcement of adequate size, correctly locatedb) Inclusion of pozzolan to reduce temperature risec) Provision of adequate contraction jointsd) Effective curing

3. Leaching and liquid penetration due to:a) Flowing liquid§

b) Pondingc) Hydraulic pressure

3. Structural designa) Minimize areas of contact and turbulenceb) Provision of membranes and protective-barrier system(s)||

c) Provision of adequate drainage and through-flow*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.‡Poor curing procedures result in flaws and cracks.#Resistance to cracking depends on strength and strain capacity.§Movement of water-carrying deleterious substances increases reactions that depend on both quantity and velocity of flow.||Concrete that will be frequently exposed to chemicals known to produce rapid deterioration should be protected with a chemically-resistant protective-barrier system.



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 207.1R, ACI 207.2R, and ACI 224R apply.

Conductive coatings applied at the time of construction as part of a cathodic protection system may provide addi-tional protection for concrete exposed to saline groundwater. Coatings that significantly restrict the evaporation of free water from the interior of concrete can reduce resistance to freezing and thawing.

7.3—Acid attackIn general, portland-cement concrete does not have good

resistance to acids, although some weak acids can be toler-ated, 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. Sewage that is not sufficiently aerated can form sulfuric acid (Flemming 1995; Sydney et al. 1996). Acids may occur in runoff from some mines and in some industrial waters. Peat soils, clay soils, and alum shales may contain sulfide-bearing minerals such as pyrite that produce sulfuric acid on oxidation. Further reactions can produce sulfate salts, which may lead to sulfate attack (Hagerman and Roosaar 1955; Lossing 1966; Bastiensen et al. 1957; Mourn and Rosenquist 1959). Some mineral waters may contain high concentrations of dissolved carbon dioxide and hydrogen sulfide; such solutions can be acidic and highly aggressive to concrete (RILEM 1962; Thornton 1978). Organic acids may come from farm silage or from manu-facturing 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. 1996). This can be of considerable concern in the case of floors, even where structural 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 hydra-tion product. At room temperature, portlandite decomposes at pH below 12.4, ettringite decomposes at pH below 10.4 (Warren and Reardon 1994), and C-S-H starts to decompose when pH drops to around 10 (Beaudoin and Brown 1992). 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

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 typi-cally produce a silica gel that has little binding capacity. The decomposition products resulting from the decalcification 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 lime-fly ash paste 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 (Ca2+) and bicarbonate (HCO3

–) ions. Detailed discussions of carbonate equilibria in natural systems can be found else-where (Stumm and Morgan 1995; Krauskopf and Bird 1995; Butler 1998). 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 significant. This type of attack has been referred to as aggressive CO2 attack in the literature when, in fact, it was not CO2 but carbonic acid attack. Test criteria relating to carbonic acid attack have used the term “aggressive CO2” (reported as milligrams of CO2 per liter) for what was really carbonic acid in sufficient concentration to attack concrete. Good discussions of this topic may be found in Lea (1971) and Hewlett (1998).

Waters that have potentially harmful concentrations of carbonic acid tend to have pH values ranging from approxi-mately neutral to slightly acidic. Because the rate of attack depends on both the properties of the concrete and concen-tration of the carbonic acid, neither the pH nor the amount of free CO2 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 CO2 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 CO2) can result in rapid carbonation and attack of the hydrated cement paste. However, freely moving waters with 10 mg/L or less of carbonic acid (reported as mg/L of CO2) can also result in significant carbonation (Terzaghi 1948, 1949; Hewlett 1998). DIN 4030 includes both criteria and a test method for assessing the potential of damage from carbonic-acid-bearing water.



7.3.3 Mitigation—The use of pozzolanic materials such as fly ash and silica fume will decrease the CH content and increase the resistance of concrete to acids (Sellevold and Nilson 1987). 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 appro-priate protective-barrier system.

7.4—Fresh waterFresh 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 fluorinated. Fresh waters can also occur in industrial, manufacturing, and other facilities where distilled waters are produced or used in various processes.

In nature, lightning produces weak nitrous, nitric, sulfu-rous, 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. Significant chemical attack by fresh water, however, is virtually unreported. That concrete is not significantly dete-riorated by fresh water is evidenced by highways, culverts, pipes, and buildings that are built with the full expectation that their function will not be significantly affected by such exposure during their expected service life.

7.4.2 Reaction mechanisms—Very pure waters are aggres-sive because they are undersaturated with respect to the CH, C-S-H, and calcium carbonate components of the cementi-tious paste. In general, the reaction mechanisms associ-ated with freshwater exposures are similar to those of acid attack because the attack involves preferential dissolution or leaching of soluble cement hydration products. Conse-quently, constant replenishment of fresh water may accel-erate deterioration and flowing water may be aggressive as well. Falling water from devices such as gutters and down-spouts 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 before significant carbonation. Such studies are often inapplicable to field concrete. Some fresh water that is undersaturated with respect to carbonates may attack carbon-ated concrete. Even under decades of such exposure, the dete-rioration 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 replen-ishment, and shelter against falling water are also important.

7.5—CarbonationCarbonation 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 1987). Carbonation begins at the exposed surface of concrete to form an outer layer of carbonate-bearing compounds, reducing the porosity of the surface (Parrott 1987). 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 1987). Prolonged moist curing may delay carbonation (Parrott 1987). Carbonation can be either beneficial or harmful, 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 Occurrence7.5.1.1 Carbonation of fresh concrete—The carbon-

ation 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 exces-sive surface cracking or a weak powdery residue (laitance) on the surface. Severe carbonation prior to final set can result in a less wear-resistant surface.

7.5.1.2 Carbonation of early-age concrete—Immature concrete is more susceptible to carbonation than mature concrete because its matrix has not hydrated sufficiently to limit permeability. Carbonation will, therefore, prog-ress 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. 1988).

7.5.1.3 Carbonation of mature concrete—Aged concrete is carbonated to some degree. Carbonation of concrete occurs at exposed surfaces. More-permeable, coarse, open-finished 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 mechanisms7.5.2.1 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 CO2. 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 insignificant, and above 75 percent relative humidity, mois-ture in the pores restricts CO2 penetration (Verbeck 1958).

7.5.2.2 Carbonation by carbon species in water—Carbon-ation can also take place when concrete is exposed to water containing sufficient concentrations of carbonate or bicar-



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 (CO3

2–) and bicarbonate (HCO3

–) 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 (1995), Krauskopf and Bird (1995), and Butler (1998).

7.5.2.2.1 Alkalinity of water (carbonate alkalinity buffer)—Carbonate alkalinity can play a critical role in the service life of concrete. In general, when carbonate alkalinity, which is a concentration of carbonate and bicarbonate ions, is sufficient 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 alka-linity, causing the selective dissolution of carbonate-bearing compounds of cement. This leaching attack, referred to as aggressive CO2 attack in past literature (Biczok 1967; Lea 1971; Ibrahim et al. 1997), is better termed “low-carbonate alkalinity attack”.

Where free CO2 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, low-carbonate 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 CO2 attack, is the leaching of the cementitious carbonate-bearing compounds until an equilib-rium 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 (1971), Biczok (1967), and Hewlett (1998).

7.6—Industrial chemicalsIndustrial chemicals may attack concrete in different

ways, depending on the type or classification of the chem-ical, its concentration, the duration of exposure, and inter-actions 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, flammables, explosives, cryogenic, and other process-specific conditions or combinations of these conditions. Refer to PCA (2001) and ACI 515.2R 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

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 infiltrate the surrounding concrete.

7.6.2 Reaction mechanisms—Because industrial chemi-cals range widely from highly alkaline to very acidic, a detailed discussion of the potential reaction mechanisms is beyond the scope of this guide. Biczok (1967) 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 dura-bility 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-and-thawing 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 supplemen-tary cementitious materials (SCMs) in the design mixture in conjunction with good curing. Reinforcing steel protec-tion can be achieved by several methods, including sealing the concrete, removing any surface contamination from the reinforcement, improving the bond between the reinforce-ment 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 fillers, joint sealers, waterstops, and surface sealers that are resistant to and compatible with the chemicals that the struc-ture is expected to be exposed to should be specified. Every-thing should be installed or applied in full compliance with manufacturers’ recommendations. Wet curing is the recom-mended 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 shrinkage-compensating concrete minimizes shrinkage cracking and the number of contraction joints, it reduces the potential leakage paths in an industrial environment.



7.7—Deicing and anti-icing chemicalsThe use of chemical agents has a long history in helping

maintain safe winter driving conditions. Such maintenance generally involves two different strategies: deicing and anti-icing. 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 pave-ment. 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 significant 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 (CaCl2), magnesium chloride (MgCl2), and many commer-cially available products comprising combinations of these salts. Some chloride-based deicers may contain significant concentrations of other chemicals. Other product formula-tions include MgCl2-based agricultural products. The use of MgCl2 deicers has increased significantly since 2000, particularly in the western United States. In part, this growth resulted because MgCl2 is more effective at lower tempera-tures than NaCl and CaCl2.

The introduction of nonchloride deicers stemmed from concerns of corrosion of reinforcing steel and environmental impacts on vegetation, and because they are also more effec-tive at lower temperatures than chloride-based deicers. 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 1970s. Their use has become more widespread with decreasing production costs. Urea was traditionally used for airfield pavements but is less commonly used now due to environmental concerns. Ethylene and propylene-based glycols are in widespread use, primarily for deicing and anti-icing of aircraft. They are also occasionally used for mainte-nance of airfield pavements, sometimes in combination with urea. A new generation of alkali-acetate and alkali-formate deicers and anti-icers emerged in the late 1980s and early 1990s to replace urea in deicing airfield 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 reac-tions that can lead to portlandite dissolution. The dissolu-tion of portlandite may increase the porosity of the concrete and lower the pH of the pore solution, which may desta-bilize 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 CaCl2 is aggressive to concrete (Collepardi et al. 1994). Among other

mechanisms, CaCl2 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. Chlo-ride may also accelerate alkali-silica reaction (ASR) under some conditions (Chatterji et al. 1986). MgCl2 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 significant deterioration in pavements by cracking, delamination, and, ultimately, disin-tegration. Disruptive oxychlorides have also been found in mortars exposed to MgCl2 (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 1995; Santagata and Collepardi 2000) report that exposure to calcium-magnesium acetate deicer solutions significantly degrades the cement matrix, resulting in loss of mass and compressive strength. The reaction mecha-nisms are similar to those of MgCl2 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 of some airfield concrete pavements exposed to alkali acetate and sodium formate deicers has caused concern. Research is ongoing to investigate potential links between these deicers and the 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 relatively rapid deterioration of joints in pavements and exterior flatwork. 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. 2012). The relationship between chemical deicers and accelerated joint deterioration is an area of active research and several mechanisms are proposed to explain this deterio-ration. In addition to depressing the freezing point of water, deicing and anti-icing chemicals induce fundamental changes in the physical and chemical properties of solutions that fill joints, such as their viscosity, surface tension, and sorption (Spragg et al. 2011; Villani et al. 2014a). 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. 2014b; Farnam et al. 2014). Concrete exposed to high concentra-tions of chemical deicers tends to have air voids filled with secondary deposits that include ettringite, portlandite, and oxychloride minerals (Sutter et al. 2006; Peterson et al. 2013). Oxychloride phases are commonly observed in labo-ratory-based studies, but documenting their presence in field concrete remains difficult due to their instability (Peterson et al. 2013). Ettringite deposits may accelerate the saturation of concrete (Stark and Bollmann 1999) and possibly diminish



the efficacy of air void systems, 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 2012). Some cracking has been detected in concrete exposed in the laboratory to deicer solutions at temperatures above freezing (Farnam et al. 2014). This cracking may occur as a result of mineralogical changes in the cement paste that result from reactions involving deicing solutions, although the specific reaction mechanisms 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 first year of service. Mitigation also requires the provision and maintenance of adequate drainage to mini-mize duration of exposure.

7.8—Environmental structuresEnvironmental 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 structures defined in ACI 350.2R. When the concentration of the contained chem-icals is sufficient, 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 envi-ronmental 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 chemi-cals 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 classifies 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 differenti-ates Group 3 chemicals according to the rate at which they will corrode concrete under typical exposure conditions associated with environmental structures. Group 3a chemi-cals 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.

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 corro-sion 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 sulfide gas. Aerobic bacteria present on wet sewage facility walls, such as pipe linings, may convert hydrogen sulfide into sulfuric acid. In these environments, significant deterioration of the concrete can occur from the dissolution of cementi-tious phases such as portlandite and C-S-H, the deposition of secondary products such as gypsum, and leaching of ferru-ginous 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 (MgNH4PO4·6H2O). Struvite deposits can impede the efficiency of 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 struc-tures, including specific 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 mate-rials; 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 chem-ical. Refer to ACI 350 for requirements concerning the use of protective barriers.

CHAPTER 8—PHYSICAL SALT ATTACK

8.1—IntroductionPhysical salt attack is a deterioration mechanism caused

by crystallization of salts in pores near concrete evapora-tive surfaces. The mechanism is called physical salt attack because chemical reactions between concrete and crystal-lizing salt are not involved (Mehta and Monteiro 2006). Deterioration ranges from very fine surface crumbling and scaling, which is primarily cosmetic, to severe progressive



disintegration. Deterioration occurs at evaporative surfaces above the soil line or on flatwork where water may wick (Fig. 8.1a and 8.1b). The most common salts linked to physical salt attack include, in order of decreasing aggres-siveness, sodium sulfate, sodium carbonate, and sodium chloride.

Physical salt attack has also been called salt crystalliza-tion, salt hydration distress, salt damp, and salt weathering. Reports identifying physical salt attack include Bates et al. (1913), Wig et al. (1917), Reading (1975, 1982), Novak and Colville (1989), Yen and Bright (1990), Haynes et al. (1996), Hime et al. (2001), and Erlin and Jana (2003). Physical salt attack on concrete began to receive attention in the early 1990s, with research initiated by Folliard and Sandberg (1994) and theoretical developments by Scherer (1999). Prior research generally focused on chemical sulfate attack on concrete, while physical sulfate attack was often overlooked or misidentified.

Physical salt attack also occurs as a deterioration mecha-nism on exposed stone and brick masonry, as well as other porous materials. This deterioration mechanism is an impor-tant 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 1970; Goudie and Viles 1997; Winkler 1997).

Scaling of concrete 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 dete-riorate by chemical and physical mechanisms (Chapter 4) (Mehta and Monteiro 2006; Valenza and Scherer 2007).

8.2—OccurrencePhysical salt attack occurs throughout the world. The

process requires water-soluble salts from seawater, ground-water, soil, and other sources and ambient environmental conditions, which usually involve fluctuations 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 condu-cive to physical salt attack.

Physical salt attack may occur even when salt concentra-tions 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

(atmospheric, ground, and irrigation). This involves deter-mining 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.

Fig. 8.1a—Physical salt attack on concrete residential foun-dation. Scaling of 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).

Fig. 8.1b—Physical salt attack on concrete garage slab caused by sodium carbonate. Scaling occurred along an evaporation front (Haynes and Bassuoni 2011).



8.3—BackgroundMuch of the early literature reported evidence of physical

salt attack by observations of field-exposed concrete, usually in sulfate-bearing soils and occasionally in carbonate-bearing soils. In 1912, tests were conducted by Bates et al. (1913) to investigate the cause of field 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 concentra-tions of alkali salts. Williams and Furlong (1926) observed that crystallization of salts in concrete pores occurred along with some other unknown chemical action. The U.S. Bureau of Reclamation (USBR) (1963) noted that salts such as sodium carbonate can cause surface disintegration by crys-tallizing in concrete pores, and that this action appears to be purely physical. Later, Reading (1975, 1982) reported that a concrete tailrace wall of a dam showed deterioration caused by physical action of sodium sulfate.

One of the more significant test programs identifying phys-ical salt attack was the Portland Cement Association’s long-term field tests on concrete exposed to sulfate-containing soils (McMillan et al. 1949; Stark 1982, 1989b, 2002). Overall, thousands of specimens were tested. The early work by McMillan et al. (1949) identified and discussed the anhy-drous form (thenardite [Na2SO4]) and the decahydrate form (mirabilite [Na2SO4·10H2O]) of sodium sulfate, but these salt phases were not associated with a deterioration mechanism. Stark (2002) examined deteriorated specimens microscopi-cally and identified physical salt attack as a main cause of deterioration in the field-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 (1982, 1989b, 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. 1996; Bassuoni and Nehdi 2009).

Folliard and Sandberg (1994) conducted exploratory tests on small concrete specimens to identify mechanisms of dete-rioration by inducing salt crystallization due to hydration, evaporation, and changes in temperature. The most aggres-sive 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 supersatu-rated with respect to mirabilite as the temperature decreased; hence, mirabilite precipitation or growth was likely respon-sible for deterioration. Haynes et al. (2008, 2010) reported on a 3-year study in which concrete was partially submerged in sodium sulfate, sodium carbonate, or sodium chloride solu-tion and found that cycling of ambient environmental condi-tions caused more deterioration than when ambient condi-tions 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

minor deterioration in laboratory testing in comparison to the other salts.

8.4—MechanismCrystallization pressure is the primary cause of physical salt

attack. Scherer (2004a) provides a summary of the equilib-rium thermodynamics that govern the development of crystal-lization 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 crys-tallizes, 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 recrys-tallize by a different mechanism. They can experience phase changes due to changes in ambient temperature and rela-tive humidity. Thenardite (Na2SO4), 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 mois-ture from the air and dissolve, from which the decahydrate form (mirabilite [Na2SO4·10H2O]) 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 314 percent increase in crystal size. For sodium carbonate, conver-sion of the low-hydrate phase (thermonatrite [Na2CO3·H2O]) to the decahydrate phase (natron [Na2CO3·10H2O]) is accom-panied by a 260 percent increase.

Factors important to distress by crystallization pressures are the degree of supersaturation of the solution, the distri-bution 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, efflorescence is usually not linked to salt-based deterioration because salt crystallization occurs on exposed surfaces where such growth is easily accom-modated. Evaporation fronts, however, can develop below exposed surfaces when fine 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 subfloresence rather than efflorescence. When subfloresence



occurs, damage may result because pressures develop in the fine pores, which can cause microcracking.

8.5—RecommendationsResearch directed to better understand this mechanism of

deterioration is needed, particularly in the areas of defining 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.

Specific 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 long-term study on concrete exposed to sulfate soils is the main refer-ence 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 fly 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.

CHAPTER 9—CORROSION OF METALS AND DEGRADATION OF OTHER MATERIALS

EMBEDDED IN CONCRETE

9.1—IntroductionUnderstanding 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 mech-anisms 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 fluid of the portland cement paste. The adequacy of 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 envi-ronment. A more comprehensive treatment of the subject can be found in ACI 222R and Broomfield (2007).

9.2—General principles of 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 sufficiently severe that cracking, spalling, or both, occur and eventually the structural element may not perform its intended function.

9.2.2 Protection mechanism in concrete9.2.2.1 General—As described in ACI 222R, the high

alkalinity, with a pH greater than 12.5, 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 film to form on the steel surface, termed a “passive film”. The composition of this film depends upon the metallurgy of the metal and is under-stood to be a combination of hydroxides and oxides. This film 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 film is maintained. If the passive film breaks down, termed “depassivation,” corrosion rate accelerates and the propagation phase begins. The film can break down locally so that localized corrosion results. If breakdown occurs over larger areas, more uniform general corrosion takes place. The primary causes of film breakdown include:

a) Chemical, physical, or mechanical degradation of the concrete cover

b) Chloride penetration to the reinforcementc) Carbonation of the concrete to reinforcement depthd) 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.

2H2O + O2 + 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 expan-sion 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 insufficient oxygen supply—The passive layer requires an oxygen flux corresponding to approximately 0.2 to 0.3 mA/m2 (1.3 × 10–4 to 1.9 × 10–4 mA/in.2). If the oxygen flux is less than this, the passive film will gradually reduce in thickness, exposing bare steel. The result is corrosion at an extremely low rate (corresponding to the oxygen flux), but at an active, though very negative, potential (Fidjestøl and Nilsen 1980). Such low oxygen



diffusion can occur in submerged concrete, in seawater or fresh water (Fidjestøl et al. 1985), or by exposure below the groundwater table.

9.2.2.4 Carbonation—The reaction between CO2, commonly from the ambient air, and cement paste can reduce the pH to less than 9, resulting in the loss of passivity. Corro-sion 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 corro-sion. 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 concentra-tion necessary to cause breakdown of the film depends on the pH and composition of the pore water, the quality of the oxide film, and the characteristics of the steel 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 pore-fluid composition, chiefly due to the use of supplementary cementitious materials (SCMs) such as fly ash, slag, and silica fume, the limits for chlorides in new construction as given in ACI 222R are still conservative and appropriate. ACI 318-14 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 of conflicting reports on chloride thresholds, the effects of different exposure environments, and materials combina-tions. 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 of corrosion9.3.1 General—Once corrosion has been initiated, a struc-

ture 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 based on suffi-ciently controlling the rate of dissolution of corrosion prod-ucts 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

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 flux from the concrete surface to the steel.

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 sufficiently large, however, most of the potential differ-ence is spent in overcoming the voltage drop IR, caused by the current flow 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 corrosion can be insignificant despite large differences in half-cell potential.

9.4—Corrosion-related properties of 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 15 to 25 percent of the paste. While the pH of saturated solu-tions of CH is only 12.4, the pH of 13.5 to 14 often found in concrete pore water (Justnes and Nygaard 1994) is explained by the OH– ions associated with alkalis in the concrete.

The presence of C3A in the cement can have two benefits: reducing both chloride ingress and in binding admixed or intruding chlorides. This was first established by Verbeck (1968) and has since been confirmed by other research (Rasheeduzzafar et al. 1992). The main conclusion of this work is that the use of very low C3A (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 1991; Horiguchi et al. 1994; Branca et al. 1992), improvements in pore distribution and permeability can counteract this deple-tion in CaOH2 (Torii and Kawamura 1994; Hakkinen 1992). 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. 1994; Fidjestøl 1987, 1991; Alonso et al. 1992).



In the 1990s, the development of corrosion-resistant concrete focused on using blends of portland cement and other cementitious materials (Baweja et al. 1994; Maage and Helland 1991; Maage et al. 1994; Berke et al. 1991; Collepardi et al. 1994; Anqi et al. 1991; Hussain and Rasheeduzzafar 1994; Decter et al. 1989; Haque et al. 1992; Ozyildirim 1994).

Malhotra et al. (2000) and Smith et al. (2004) all reported that concrete containing moderate to high volumes of fly ash exhibited superior resistance to the penetration of chlorides and improved corrosion resistance.

Slag cement has been used in marine work since the early 1900s; the experiences with respect to resistance against chloride-induced corrosion are generally good (Wiebenga 1984; Hope et al. 1985; Pal et al. 2002). Slag cement has also been shown to improve resistance to penetration of deicer salts (McGrath and Hooton 1997; 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 1991; Pettersson 1995; Fidjestøl 1987, 1993; Fidjestøl and Frearson 1994; Alonso et al. 1992; Berke et al. 1992; Gautefall and Vennesland 1985; Zhang and Gjørv 1991; Fidjestøl and Justnes 2002; Skjølsvold et al. 2007).

9.4.3 Aggregates—Aggregates can contain chloride salts, particularly those aggregates that have been exposed to seawater or whose natural sites are in groundwater containing chloride. Sedimentary rock formed in ancient seabeds can also contain significant amounts of chlorides. There have been reported instances (Gaynor 1985) 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 C1524 can be used to determine the water-extractable chloride content of aggre-gates that could potentially contribute to corrosion initiation.

9.4.4 Mixing water—Potable mixing water can contain small amounts of chloride. ASTM C1602/C1602M has optional limits for chlorides in mixing water: 500 ppm for prestressed concrete, bridge decks, or as otherwise desig-nated; and 1000 ppm for other reinforced concrete in moist environments.

9.4.5 Admixtures9.4.5.1 General—Admixtures containing significant

concentrations of CaCl2 should not be used in concrete containing embedded metal. Some water-reducing admix-tures can contain chloride to improve admixture perfor-mance, 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 212.3R.

9.4.5.2 Accelerators—Accelerating admixtures, other than those based on CaCl2, 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 accel-erators are calcium formate, sodium thiocyanate, calcium nitrate, and calcium nitrite. It is generally accepted that formates (Holm 1987) 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 corrosion-inhibiting admixtures are common commercially: 1) amine carboxylate; 2) amine-ester organic emulsion; 3) calcium nitrite; and 4) an organic alkenyl dicarboxylic acid salt (ACI 212.3R). 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 film at the steel surface (Nmai et al. 1992; Bobrowski and Youn 1993). Laboratory evaluations indicate that amine-ester organic emulsion will delay the onset and reduce the rate of corrosion (Nmai et al. 1992; Nmai and Krauss 1994).

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 1985; Berke and Roberts 1989). The ratio of chlo-ride ions to nitrite ions is important. Studies (Berke 1987) 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./yd3 (10 to 30 L/m3) of concrete is the common range. Berke and Rosenberg (1989) 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. (1994). 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 corrosion9.5.1 General—In mitigating corrosion, these critical

points should be evaluated in terms of concrete service life:a) Initiation of corrosionb) Corrosion products becoming visible (staining)c) Serviceability, including cracking, spalling, or bothd) The load-carrying capacity of the structure if it is seri-

ously 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 appro-priate in design.

The corrosion of steel in concrete is dependent on the envi-ronment in which the structure is exposed. Table 19.3.1.1



in ACI 318-14 specifies exposure categories for corrosion 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. Corro-sion 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 365.1R. 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 process9.5.2.1 Concrete quality and cover over steel9.5.2.1.1 Cover depth—Extensive studies (Clear 1976;

Pfeifer et al. 1987; Marusin and Pfeifer 1985) 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. (12 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 moderate-to-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 influence of cover depths and recom-mended cover depths for bridges of 2.75 in. (70 mm). ACI 318-14 specifies concrete cover requirements for Exposure 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 steel performs significantly 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. 1987). Simi-larly, the proper use of SCMs can extend the time to corro-sion and reduce its rate.

ACI 318 and other specifications place strict requirements on the mixture proportions for severe chloride environments. The Norwegian Public Roads Administration (2009) has specific 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 specifications for the Great Belt (Storbælt) project required a low w/cm with the use of silica fume and fly ash to provide a 100-year service

life (Storebælt Technical Publications 1999). 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 (10 km) connection across Øresund from Sweden to Denmark (Henriksen et al. 2000).

Admixed chlorides can also influence 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 specifications in other countries, such as CSA A23.1-14/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 conven-tional reinforcing steel, and often suffer from additional corrosion mechanisms (9.6). The effect of corrosion mecha-nisms and high stresses in these steels is to make extra restric-tions necessary compared to black steel. Structure type and importance, fabrication methods, and exposure conditions also influence corrosion performance of these steels.

For post-tensioned structures, the Federal Highway Administration (FHWA) (2012) 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 infiltration rate than diffusion processes, and can establish chloride concentration cells that accelerate corrosion. ACI 318-99 concluded that the role of cracks in the corrosion of reinforcement is controversial and, without sound scientific data relating crack width to corrosion activity, modified 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 318 does not provide explicit limits for crack widths in the equation, this equation does indirectly limit the maximum crack width to between 0.016 and 0.02 in. (0.4 and 0.52 mm). In addition, the presence of cracking effects on corrosion

Table 9.5.2.1.2—Limits to chlorides in newly constructed concrete (as recommended by ACI 222R)

Acid-soluble(ASTM C1152/

C1152M)

Water-soluble(ASTM C1218/

C1218M)Soxhletmethod*

Prestressed concrete 0.08 0.06 0.06

Reinforced concrete in wet

conditions0.10 0.08 0.08

Reinforced concrete in dry

or protected conditions

0.20 0.15 0.15

*Soxhlet method described in ACI 222.1R.Note: All chloride contents expressed as percent Cl– by mass of cement.



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 corro-sion, 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 (magne-sium hydroxide and calcium carbonate) within a fairly short period (Espelid and Fidjestøl 1986; Buenfeld and Newman 1986). ACI 224R and ACI 224.1R 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 of reinforcement, sufficient 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 modi-fication 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 CaCl2 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 1982). Main-taining an internal relative humidity below 50 percent, however, is not always possible.

While the surface regions of exposed concrete struc-tures will have high or low electrical conductivity values depending on the wetting and drying conditions of the envi-ronment, the concrete interior usually requires long drying periods to achieve low electrical conductivity. Pfeifer et al. (1987) found that 7 to 9 in. (180 to 230 mm) thick rein-forced 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, fly ash, and, in particular, silica fume, will give concrete resistivi-ties far in excess of what is provided by portland-cement concrete (Cabrera and Ghoddoussi 1994; Fidjestøl and Frearson 1994; Gautefall and Vennesland 1985; Berke 1988). 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-deic-cation. Field investigations after more than 20 years confirm long-term performance of SCMs (silica fume) (Fidjestøl and Justnes 2002; Skjølsvold et al. 2007).

The high electrical resistivity of blended binder systems is confirmed by tests using ASTM C1202 (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 chlo-ride diffusion determined by more conventional diffusion

or ponding tests (Detwiler and Fapohunda 1993; Wolsiefer 1991; Misra et al. 1994; Burg and Öst 1992). 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 aspects9.5.3.1 Workmanship—Good workmanship is vital for

securing uniform concrete with low penetrability. For low-slump concrete, segregation and honeycombing can be avoided by good consolidation. Meeting the requirements of the specifications pertaining to durability are essential.

9.5.3.2 Reinforcement detailing—Two factors are impor-tant 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 (10 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 15 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 specified. Limiting early thermal stresses is also important. Good curing reduces transport parameters as well as charge passed in ASTM C1202 (or AASHTO T277) (Acker et al. 1986; Marusin 1989). In addition, due to insufficient 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 essen-tial. Properly supported screeding equipment and correct supports for the reinforcement are important for attaining the cover protection specified. The use of 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. 1993).

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 of 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: field inspection of concrete bridges (Fluge and Blankvoll 1995) 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 struc-ture than on the windward side. Similarly, moving bridge columns away from traffic splash will reduce the chloride exposure of the concrete.



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 1994), 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 515.2R for a comprehen-sive 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), latex-modified concrete overlays (Clear and Hay 1973; FHWA 1975), concrete containing silica fume, and concrete containing high-range water-reducing admixtures

b) Epoxy-coated reinforcing steelc) Corrosion-resistant steels (Rasheeduzzafar et al.

1992; Trejo and Pillai 2004; Clemeña and Virmani 2004; Williamson et al. 2003)

d) Waterproof membranes (Van Til et al. 1976)e) Surface protective-barrier systems produced from

silanes, siloxanes, epoxies, polyurethanes, and methacry-lates (Van Daveer and Sheret 1975)

f) Cathodic protection

9.6—Corrosion of prestressed steel reinforcementThe 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 embrittle-ment 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 sulfide, 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 reinforce-ment 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 afflicted element(s), and are difficult to locate via inspection or testing when the structure is in service. Preventive measures

during fabrication/construction are needed to reduce the risk of occurrence. ACI 222.2R provides addition details on the prestressing systems, deterioration, protection, field 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 finished 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 ware-house, 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 rein-forcement 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 water-tight, 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 there-after to avoid exposure. Systems that employ grease canis-ters at end anchorages should be periodically inspected to confirm 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 of materials other than steel9.7.1 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, particu-larly 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 1984; Erlin 2006).



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 1968; 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 1989).

9.7.3 Lead—Lead in damp concrete will be attacked by the CaOH2 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). Corro-sion of copper pipes, however, has been reported where ammonia is present. Also, there have been reports that small amounts of ammonia, and 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 inten-tionally embedded in concrete. Available data are conflicting as to the benefit, if any, of this coating (Cook 1980; Stark and Perenchio 1975; Hill et al. 1976; Griffin 1969; Federal Highway Administration 1976). A chromate dip on the galvanized bars or the use of 400 ppm of chromate in the mixing water is recommended to prevent hydrogen evolu-tion 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. Corro-sion inhibitors, such as calcium nitrite, have been shown to improve the corrosion resistance of zinc in concrete (Berke and Rosenberg 1989; Page et al. 1989).

9.7.6 Other metals—Chromium and nickel alloys gener-ally 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 steel may be economically justified 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 140°F (60°C) and chlo-ride 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, reinforce-ment, 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, specific test data should 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 fibers in the concrete and the embedment of timber (Erlin 2006).

The use of untreated sawdust, wood chips, or fibers usually results in slow-setting and low-strength concrete (Erlin 2006). The addition of hydrated lime equal to one-third to one-half the volume of the cement is usually effec-tive in minimizing these problems. The further addition of up to 5 percent of CaCl2 dihydrate by mass of cement has also helped to minimize these problems (Erlin 2006). CaCl2 in such amounts can cause corrosion of embedded metals, however, and can have adverse effects on the concrete.

Another problem with such concrete 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—SummaryPortland-cement concrete can provide excellent corrosion

protection to embedded steel. When corrosion occurs, the costs of repairs can be exceedingly high. The use of high-quality 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 environ-ments and remedial measures where corrosion is occurring.

CHAPTER 10—ABRASION

10.1—IntroductionThe abrasion resistance of concrete is defined as the ability

of a surface to resist being worn away by rubbing and friction. Abrasion of floors and pavements can result from production operations or foot or vehicular traffic. Abrasion resistance



is, therefore, of concern in industrial floors. Wind, water, or waterborne particles can also abrade concrete surfaces (Price 1947). 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 briefly in this guide; the subject is treated in more detail in ACI 210R and 210.1R.

10.2—Testing concrete for resistance to abrasionResearch to develop meaningful laboratory tests on

concrete abrasion has been ongoing since the early 1900s. There are several different types of abrasion, and no single test method has been found that is adequate for all condi-tions. A detailed description of abrasion/erosion test methods can be found in Bakke (2006). Prior (1966) described four broad areas related to abrasion:

1. Wear on floor and slab construction; Table 10.2 shows classes of floor traffic and use, and the special considerations required for good wear resistance (ACI 302.1R).

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, spill-ways, tunnels, bridge piers, and abutments, due to the action of abrasive materials carried by flowing liquid (attrition and scraping).

4. Cavitation action on concrete in dams, spillways, tunnels, and other hydraulic structures due to high flow velocities and negative pressures.

ASTM C779/C779M covers three operational proce-dures for evaluating floor 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 (1994) commented that the most repro-ducible 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 field-testing sites should be made on the basis of providing representative results.

To set limits for abrasion resistance of concrete, it is neces-sary 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 C1138 uses agita-tion of steel balls in water to determine abrasion resistance.

ASTM C418 uses a sandblasting apparatus to measure the depth or wear to simulate comparative sand-impinged wear resistance. This test provides a means for evaluating resis-tance to abrasion caused by wind-blown sand.

The abrasion resistance of pervious concrete can be measured using ASTM C1747/C1747M. In this test method, the impact and abrasion resistance of the pervious concrete is determined from the percent mass loss from three cylin-drical 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 deflection 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. 2011).

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 of these tests target specific exposures or structural types (for example, railroad ties). Application of one or more of these tests to assess future performance of a specific 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 of concrete

The abrasion resistance of concrete is a progressive phenomenon. Initially, the resistance is related to compres-sive strength of the wearing surface. Therefore, initial judg-ments regarding relative floor wear can be made on the basis of compressive strength.

As softer paste wears away, however, the particles of fine 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 field experience have generally shown that abra-sion resistance is proportional to the compressive strength of concrete (Scripture et al. 1953; Witte and Backstrom 1951). Because abrasion occurs at the surface, it is critical that the surface strength be maximized. Resistance can be greatly improved by the use of dry shakes and toppings, finishing techniques (10.4.4), 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 finishing of floor slabs to obtain an abrasion-resistant surface by encouraging the use of both power-trowel finishing and adequate curing (Kettle and Sadegzadeh 1987).



With a given concrete mixture, compressive strength at the surface is improved by:

a) Avoiding segregationb) Eliminating bleedingc) Properly timed finishing

d) Minimizing surface w/cm (water addition to the surface during finishing 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 302.1R)

Table 10.2—Floor classifications* and considerations to improve abrasion resistanceClass Anticipated traffic type Use Special considerations Final finish

1. Exposed Exposed surface—foot traffic Offices, churches, multiunit residential, decorative

Uniform finish, nonslip aggregate in specific areas, curing

Colored mineral aggregate, color pigment or exposed aggregate, stamped or inlaid patterns, artistic joint layout, curing, surface treat-ment, maintenance

Normal steel-troweled finish, nonslip finish where required

Burnishing or polishing to enhance sheen as required

2. Covered Covered surface—foot traffic Offices, churches, commercial, multiunit residential, institu-tional with floor coverings

Flat and level slabs suitably dry for applied coverings, curing

Light steel-troweled finish

3. Topping Exposed or covered surface—foot traffic

Unbonded or bonded topping over base slab for commercial or non-industrial buildings where construction type or schedule dictates

Base slab—good uniform level 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. (19 mm) minimum thickness curing

Base slab—troweled finish under unbonded topping; clean, textured surface under bonded topping

Topping—for exposed surface, normal steel-troweled finish; for covered surface, light steel-troweled finish

4. Institutional/commercial

Exposed or covered surface—foot and light vehicular traffic

Institutional or commercial Level and flat slab suitable for applied coverings, nonslip aggregate for specific areas, curing; coordi-nate joints with applied coverings

Normal steel-troweled finish

5. Industrial Exposed surface—industrial vehicular traffic such as pneu-matic wheels and moderately soft solid wheels

Industrial floors for manu-facturing, processing, and warehousing

Good uniform subgrade, joint layout, joint load transfer, abrasion resistance, curing

Hard steel-troweled finish

6. Heavy industrial

Exposed surface—heavy-duty industrial vehicular traffic such as hard wheels and heavy wheel loads

Industrial floors subject to heavy traffic; can be subject to impact loads

Good uniform subgrade, joint layout, joint load transfer required, abrasion resistance, curing

Special metallic or mineral aggregate surface hardener; repeated hard steel-troweling

7. Heavy industrial topping

Exposed surface—heavy-duty industrial vehicular traffic such as hard wheels and heavy wheel loads

Bonded two-course floors subject to heavy traffic and impact

Base slab—good uniform subgrade, reinforcement, joint layout, level surface, curing

Topping—composed of well-graded all-mineral or all-metallic aggregate. Minimum thickness 3/4 in. (19 mm)

Mineral or metallic aggregate surface hardener applied to high-strength plain topping to toughen, curing

Clean, textured base slab surface suitable for subse-quent bonded topping. Special power floats for topping are optional, hard steel-troweled finish

8.Commercial/industrial topping

As in Classes 4, 5, or 6 Unbonded topping—on new or old floors where construction sequence or schedule dictates

Bondbreaker on base slab, minimum thickness 4 in. (100 mm), abrasion resistance, curing

As in Classes 4, 5, or 6

9.Critical surface profile

Exposed surface—superflat or critical surface tolerance required; special materials-handling vehicles or robotics requiring specific tolerances

Narrow-aisle, high-bay ware-houses; television studios, ice rinks, or gymnasiums (ACI 360R)

Varying concrete quality require-ments. Special application proce-dures and strict attention to detail are recommended when shake-on hardeners are used. FF 50 to FF 125, superflat floor, curing

Strictly following techniques as indicated in 8.9 of ACI 302.1R

*Taken from Table 4.1 of ACI 302.1R-15.



f) Proper curing proceduresEconomical 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 fly ash are used in concrete, the abrasion resistance is gener-ally related to the compressive strength developed (Keshari 2009; Naik et al. 1995, 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 1987), epoxy concrete (Mirza et al. 1990), calcium aluminate cement (Scrivener et al. 1999), 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. 1953; Smith 1958). The service life of some concrete slabs, such as warehouse floors that are subjected to abrasion by steel or hard rubber-wheeled traffic, is greatly lengthened by the use of hard, tough aggregates.

The abrasion resistance of lightweight concrete is a func-tion of the concrete compressive strength; however, the use of only lightweight aggregates may not be advisable for structures with high abrasion resistance requirements (ACI 213R). 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. 2012; de Brito 2010).

Abrasion-resistant aggregates can be used either with the dry-shake method (ACI 302.1R) or as part of a high-strength topping mixture. If abrasion is the principal concern, addi-tion of high-quality quartz, traprock, or emery aggregates properly proportioned with cement will increase the abra-sion 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 abra-sion resistance and increase service life. Another advantage of using metallic aggregate is improved impact resistance, especially at joints.

The use of two-course floors using a high-strength topping is generally limited to floors where both abrasion and impact resistance are required. While providing excellent abrasion resistance, a two-course floor 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 of satisfactory floor surfaces is curing (Liu 1994; ACI 302.1R; ACI 308R). Because the uppermost part of the concrete surface is the region that is abraded by traffic, maximum strength and toughness are the most important elements for ensuring resistance to surface abrasion. This is partially accomplished through proper finishing operations, troweling techniques, and adequate and timely curing practices (10.4.4). The effect of curing efficiency (absorptivity) at the top-wearing surface has been shown to be directly related to abrasion resistance.

Curing has less effect on the abrasion resistance of deeper sections of the same concrete (Senbetta and Scholer 1984).

10.4—Recommendations for obtaining abrasion-resistant concrete surfaces

10.4.1 Factors affecting abrasion resistance—The following factors directly impact concrete strength and, therefore, abrasion resistance (ACI 302.1R):

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 of finishing operations that avoid the addition of water during troweling, and vacuum dewatering.

b) Well-graded fine 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 delami-nation if finishing 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 floors—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. (12.5 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 finishing procedures—Floating and trow-eling 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 finishing 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 finishing operations is provided in ACI 302.1R.

10.4.5 Vacuum dewatering—Vacuum dewatering is a method for removing water from concrete immediately after placement (New Zealand Portland Cement Association 1975). While this permits a reduction in w/cm, the quality of the finished surface is still highly dependent on the timing of finishing 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 recommenda-tions should be followed. Additional guidance is provided in ACI 302.1R.

10.4.7 Proper curing procedures—For most concrete floors, water curing (keeping the concrete continuously wet)



is the most effective method of producing a hard, dense surface (Shurpali et al. 2012). 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 exces-sive 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-modified 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 first 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 require-ments of ASTM C309 or ASTM C1315. They should be applied in a uniform coat immediately after concrete finishing and in accordance with the manufacturer’s recom-mendations. Recommended coverage rates will vary depending on the surface texture of the finished surface. A smoother floor 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 if the floor is subjected to traffic before curing is complete. Curing compounds should not be required for surfaces that receive paint or floor 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 (CO2), such as finishing machines, vehicles, and welding machines, should not be used without attention to proper ventilation. CO2 can adversely affect fresh concrete surfaces between the time of placement and applica-tion of a curing compound through a mechanism referred to as carbonation. The severity of the effect is dependent on the concentration of CO2, the humidity and ambient temperature, and the length of exposure to the air (Kauer and Freeman 1955; Matsuzawa et al. 2010). Early carbon-ation 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 concreteAbrasive 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 100 in each tire. One laboratory study showed that studded tires running on surfaces to which sand and salt were applied caused 100 times as much wear as tires without studs (Krukar and Cook 1973). 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 12 mm) deep can form in a single winter in regions where approximately 30 percent of passenger cars are equipped with studded tires and traffic is heavy (Smith and Schonfeld 1970). More severe wear occurs where vehicles stop, start, or turn (Keyser 1971).

Investigations have been made, principally in Scandan-avia, Canada, and the United States, to examine the prop-erties of existing concrete as related to studded tire wear (Smith and Schonfeld 1971; Keyser 1971; Preus 1973; Wehner 1966; Thurmann 1969). In 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 beneficial in resisting abrasion.

Another investigation was aimed at developing more wear-resistant types of concrete overlays (Preus 1973). Polymer cement concrete and polymer fly-ash concrete provide better resistance to wear, although at the sacrifice of skid resistance. Steel fiber concrete overlays were also tested and showed reduced wear. Exposed fibers 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 1996) summarizes available data on pavement wear and on the performance and winter accident records for studded tire use.

10.6—Skid resistance of pavementsThe 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 of fine aggregateThe microtexture is more important at speeds of less than

approximately 50 mph (80 km/h) (Kummer and Meyer 1967; Murphy 1975; Wilk 1978). At speeds greater than 50



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 1981). In time, rubber-tire traffic abrades the surface paste, which removes the beneficial macrotexture and exposes the coarse- and fine-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) of the fine and coarse aggregate.

Fine aggregates containing significant amounts of silicate minerals in the larger particle sizes will assist in slowing down surface wear and maintaining the microtexture neces-sary for satisfactory skid resistance at slow speed (Fowler and Rached 2012; Rado 2009). Certain rock types, however, polish under rubber-tire wear. These include very fine-textured limestone, dolomite, and serpentine; the finer the texture, the more rapid the polishing. Where both the fine 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 hydro-planing. An example of constructing macrotexture in pave-ment 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 finished and cured has the required durability and abrasion resistance (Ong and Fwa 2008).

10.7—ErosionConcrete erosion is the progressive removal of mass

from the concrete surface from chemical attack, abrasion, or cavitation (ACI 210R). Chemical attack erosion occurs when components of the concrete paste or aggregate are leached or dissolved. The rate of degradation can be signifi-cantly enhanced by flowing liquid due to the increased 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 pres-sure in a hydraulic system drops below the liquid vaporiza-tion 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 210R).

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 210R. Erosion by chemical attack is covered in more detail in Chapter 6 of this guide.

10.7.1 Abrasion—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 experi-ence abrasion/erosion damage (ACI 210R; U.S. Bureau of Reclamation 1997). Bridge abutments and other structures placed in rivers or other flowing bodies are also 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 aggre-gates. 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 1969; Scrivener et al. 1999; U.S. Bureau of Reclamation 1997).

10.7.2 Cavitation—The most efficient 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 flow rate, or by aeration (Frizell and Mefford 1991). 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 struc-ture 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-modified concrete has also been shown to increase 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 flow characteristics. When these methods are impractical, the use of higher-quality, erosion-resistant repair materials may prolong the service life of the struc-ture. 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 1997).



CHAPTER 11—SUMMARYThe 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 influence the durability of concrete, considering the particular mecha-nisms 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. Specifically, this guide discusses the importance of concrete’s resistance to fluid ingress as it influences 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 12—REFERENCESACI committee documents and documents published by

other organizations are listed first by document number, full title, and year of publication followed by authored docu-ments listed alphabetically.

American Concrete InstituteACI 207.1R-05(12)—Guide to Mass ConcreteACI 207.2R-07—Report on Thermal and Volume Change

Effects on Cracking of Mass ConcreteACI 210R-93(08)—Erosion of Concrete in Hydraulic

StructuresACI 210.1R-94—Compendium of Case Histories

on Repair of Erosion-Damaged Concrete in Hydraulic Structures

ACI 212.3R-10—Report on Chemical Admixtures for Concrete

ACI 213R-14—Guide for Structural Lightweight-Aggre-gate Concrete

ACI 216.1-14—Code Requirements for Determining Fire Resistance of Concrete and Masonry Construction Assemblies

ACI 221.1R-98(08)—Report on Alkali-Aggregate ReactivityACI 222R-01(10)—Protection of Metals in Concrete

against CorrosionACI 222.1R-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-14—Report on Corrosion of Prestressing Steels

ACI 222.3R-11—Guide to Design and Construction Prac-tices to Mitigate Corrosion of Reinforcement in Concrete Structures

ACI 223R-10—Guide for the Use of Shrinkage-Compen-sating Concrete

ACI 224R-01(08)—Control of Cracking in Concrete Structures

ACI 224.1R-07—Causes, Evaluation, and Repair of Cracks in Concrete Structures

ACI 301-16—Specifications for Structural ConcreteACI 302.1R-15—Guide for Concrete Floor and Slab

ConstructionACI 304R-00(09)—Guide for Measuring, Mixing, Trans-

porting, and Placing ConcreteACI 305R-10—Guide to Hot Weather ConcretingACI 306R-16—Guide to Cold Weather ConcretingACI 308R-16—Guide to External Curing of ConcreteACI 309R-05—Guide for Consolidation of ConcreteACI 318-99—Building Code Requirements for Structural

Concrete and CommentaryACI 318-14—Building Code Requirements for Structural

Concrete and CommentaryACI 350-06—Code Requirements for Environmental

Engineering Concrete Structures and CommentaryACI 350.1-10—Specification for Tightness Testing of

Environmental Engineering Concrete Containment Struc-tures and Commentary

ACI 350.2R-04—Concrete Structures for Containment of Hazardous Materials

ACI 357.1R-91(97)—Report on Offshore Concrete Struc-tures for the Arctic

ACI 360R-10—Guide to Design of Slabs-on-GroundACI 365.1R-00—Service-Life PredictionACI 515.2R-13—Guide to Selecting Protective Treat-

ments for Concrete

ASTM InternationalASTM C33/C33M-16—Standard Specification for

Concrete AggregatesASTM C109/C109M-16—Standard Test Method for

Compressive Strength of Hydraulic Cement Mortars (Using 2-in. or [50-mm] Cube Specimens)

ASTM C114-15—Standard Test Methods for Chemical Analysis of Hydraulic Cement

ASTM C150/C150M-16—Standard Specification for Portland Cement

ASTM C227-10—Standard Test Method for Potential Alkali Reactivity of Cement-Aggregate Combinations (Mortar Bar Method)

ASTM C260/C260M-10a(2016)—Standard Specification for Air-Entraining Admixtures for Concrete

ASTM C265-08—Standard Test Method for Water-Extractable Sulfate in Hydrated Hydraulic Cement Mortar

ASTM C289-07—Standard Test Method for Potential Alkali-Silica Reactivity of Aggregates (Chemical Method) (withdrawn 2016)

ASTM C295/C295M-12—Standard Guide for Petro-graphic Examination of Aggregates for Concrete

ASTM C309-11—Standard Specification for Liquid Membrane-Forming Compounds for Curing Concrete

ASTM C418-12—Standard Test Method for Abrasion Resistance of Concrete by Sandblasting

ASTM C441/C441M-11—Standard Test Method for Effectiveness of Pozzolans or Ground Blast-Furnace Slag



in Preventing Excessive Expansion of Concrete Due to the Alkali-Silica Reaction

ASTM C452/C452M-15—Standard Test Method for Potential Expansion of Portland-Cement Mortars Exposed to Sulfate

ASTM C457/C457M-12—Standard Test Method for Microscopical Determination of Parameters of the Air-Void System in Hardened Concrete

ASTM C494/C494M-16—Standard Specification for Chemical Admixtures for Concrete

ASTM C563-16—Standard Test Method for Approxima-tion of Optimum SO3 in Hydraulic Cement Using Compres-sive Strength

ASTM C586-11—Standard Test Method for Potential Alkali Reactivity of Carbonate Rocks for Concrete Aggre-gates (Rock-Cylinder Method)

ASTM C595/C595M-15—Standard Specification for Blended Hydraulic Cements

ASTM C618-15—Standard Specification for Coal Fly Ash and Raw or Calcined Natural Pozzolan for Use in Concrete

ASTM C642-13—Standard Test Method for Density, Absorption, and Voids in Hardened Concrete

ASTM C666/C666M-15—Standard Test Method for Resistance of Concrete to Rapid Freezing and Thawing

ASTM C672/C672M-12—Standard Test Method for Scaling Resistance of Concrete Surfaces Exposed to Deicing Chemicals

ASTM C779/C779M-12—Standard Test Method for Abrasion Resistance of Horizontal Concrete Surfaces

ASTM C845/C845M-12—Standard Specification for Expansive Hydraulic Cement

ASTM C856-14—Standard Practice for Petrographic Examination of Hardened Concrete

ASTM C944/C944M-12—Standard Test Method for Abrasion Resistance of Concrete or Mortar Surfaces by the Rotating-Cutter Method

ASTM C989/C989M-14—Standard Specification for Slag Cement for Use in Concrete and Mortars

ASTM C1012/C1012M-15—Standard Test Method for Length Change of Hydraulic-Cement Mortars Exposed to a Sulfate Solution

ASTM C1017/C1017M-13—Standard Specification for Chemical Admixtures for Use in Producing Flowing Concrete

ASTM C1038/C1038M-14—Standard Test Method for Expansion of Hydraulic Cement Mortar Bars Stored in Water

ASTM C1105-08—Standard Test Method for Length Change of Concrete Due to Alkali-Carbonate Rock Reaction

ASTM C1138-97—Standard Test Method for Abrasion Resistance of Concrete (Underwater Method)

ASTM C1152/C1152M-04(2012)—Standard Test Method for Acid-Soluble Chloride in Mortar and Concrete

ASTM C1157/C1157M-11—Standard Performance Spec-ification for Hydraulic Cement

ASTM C1202-12—Standard Test Method for Electrical Indication of Concrete’s Ability to Resist Chloride Ion Penetration

ASTM C1218/C1218M-15—Standard Test Method for Water-Soluble Chloride in Mortar and Concrete

ASTM C1240-15—Standard Specification for Silica Fume Used in Cementitious Mixtures

ASTM C1260-14—Standard Test Method for Potential Alkali Reactivity of Aggregates (Mortar-Bar Method)

ASTM C1293-08(2015)—Standard Test Method for Determination of Length Change of Concrete Due to Alkali-Silica Reaction

ASTM C1315-11—Standard Specification for Liquid Membrane-Forming Compounds Having Special Properties for Curing and Sealing Concrete

ASTM C1524-02(2010)—Standard Test Method for Water-Extractable Chloride in Aggregate (Soxhlet Method)

ASTM C1543-10—Standard Test Method for Determining the Penetration of Chloride Ion into Concrete by Ponding

ASTM C1556-11—Standard Test Method for Deter-mining the Apparent Chloride Diffusion Coefficient of Cementitious Mixtures by Bulk Diffusion

ASTM C1567-13—Standard Test Method for Deter-mining the Potential Alkali-Silica Reactivity of Combina-tions of Cementitious Materials and Aggregate (Accelerated Mortar-Bar Method)

ASTM C1580-15—Standard Test Method for Water-Soluble Sulfate in Soil

ASTM C1585-13—Standard Test Method for Measure-ment of Rate of Absorption of Water by Hydraulic-Cement Concretes

ASTM C1602/C1602M-12—Standard Specification for Mixing Water Used in the Production of Hydraulic Cement Concrete

ASTM C1747/C1747M-13—Standard Test Method for Determining Potential Resistance to Degradation of Pervious Concrete by Impact and Abrasion

ASTM C1778-16—Standard Guide for Reducing the Risk of Deleterious Alkali-Aggregate Reaction in Concrete

American Association of State and Highway Transportation Officials (AASHTO)

PP065-11—Standard Practice for Determining the Reac-tivity 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

T277-15—Standard Method of Test for Electrical Indica-tion of Concrete’s Ability to Resist Chloride Ion Penetration

CSA GroupCAN/CSA A3000-13—Cementitious Materials CompendiumCSA A23.1-14/CSA A23.2-14—Concrete Materials and

Methods of Concrete Construction/Test Methods and Stan-dard Practices for Concrete

CAN/CSACSA A23.2-14A-14—Potential Expansivity of Aggre-

gates; Procedure for Length Change Due to Alkali-Aggre-gate Reaction in Concrete Prisms

CSA A23.2-26A-14—Determination of Potential Alkali-Carbonate Reactivity of Quarried Carbonate Rocks by Chemical Composition



CSA A23.2-27A-14—Standard Practice to Identify Poten-tial for Alkali-Reactivity of Aggregates and Measures to Avoid Deleterious Expansion in Concrete

CSA A23.2-28A-14—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 of StandardizationDIN 4030:2006-06—Assessment of Water, Soil, and

Gases for Their Aggressiveness to Concrete – Part 1: Prin-ciples and Limiting Values

U.S. Army Corps of EngineersCRD-C 662:2010—Determining the Potential Alkali-

Silica Reactivity of Combinations of Cementitious Mate-rials, Lithium Nitrate Admixture and Aggregate (Acceler-ated Mortar-Bar Method)

EM 1110-2-2000:1994—Standard Practice for Concrete for Civil Works Structures

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