Establishing and Satisfying Thermal Requirements for Drilled Shaft Concrete Based on Durability Considerations

Abstract

Some U.S. transportation agencies have recently applied mass concrete provisions to drilled shafts, imposing limits on maximum temperatures and maximum temperature differentials. On one hand, temperatures commonly observed in large-diameter drilled shafts have been observed to cause delayed ettringite formation (DEF) and thermal cracking in above-ground concrete elements. On the other, the reinforcement and confinement unique to drilled shafts should provide resistance to thermal cracking, and the provisions that have been applied are based on dated practices for above-ground concrete. This paper establishes a rational procedure for design of drilled shafts for durability requirements in response to hydration temperatures, which addresses both DEF and thermal cracking. DEF is addressed through maximum temperature differential limitations that are based on concrete mix design parameters. Thermal cracking is addressed through calculations that explicitly consider the thermo-mechanical response of concrete for predicted temperatures. Results from application of the procedure indicate consideration of DEF and thermal cracking potential for drilled shafts is prudent, but provisions that have been applied to date are overly restrictive in many circumstances, particularly the commonly adopted 35°F maximum temperature differential provision.

Advancements in drilling and concrete technology and increased load demand for bridge foundations have fueled a recent tendency toward larger and larger diameter drilled shafts for support of transportation structures. Concerns over the potential for thermal damage in large concrete elements have led some agencies to impose so-called “mass concrete” provisions on drilled shafts, while other agencies have explicitly exempted drilled shafts from some or all requirements in mass concrete provisions. Opponents of designating drilled shafts as mass concrete correctly point to the lack of recorded evidence of thermal damage to drilled shafts, and to the unique benefits of confinement provided by the ground and heavy reinforcement concentrated near the shaft edges. Proponents of designating drilled shafts as mass concrete correctly note that no one has ever looked for thermal damage, and note the frequent observation of high temperatures in drilled shafts that have been associated with thermal damage in above-ground concrete elements. The objective of this paper, developed out of research for the U.S. Federal Highway Administration (FHWA), is to present rational procedures for establishing thermal requirements for drilled shafts, and to do so in a manner that: (a) is based on research developments in relation to mass concrete and (b) includes consideration of the unique characteristics of drilled shafts compared with above-ground concrete elements. Both aspects of this paper’s consideration of thermal requirements for drilled shafts are in contrast to the approach adopted by some agencies of applying mass concrete provisions developed at least 50 years ago for above-ground applications. Procedures described in this paper are intended to appropriately consider and address thermal requirements without unnecessarily imposing restrictive measures that complicate construction and introduce additional concrete integrity risks.

The FHWA study included a review of tremie concrete practices for the 13 state transportation agencies shown in Figure 1. The review was completed in 2015. Among the 13 agencies, practices varied in relation to whether drilled shafts were subject to mass concrete requirements. Standard specifications for all but one of the agencies include sections in relation to mass concrete, but only four of the agencies had applied provisions for mass concrete to drilled shafts: Caltrans, Florida DOT (FDOT), Minnesota DOT (MNDOT), and South Carolina DOT (SCDOT). The adopting agencies indicated that application of mass concrete provisions to drilled shafts has occurred recently and that specifications and practices were still being refined. The other nine agencies did not currently apply mass concrete provisions to drilled shafts, with some opposed to the idea for constructability and cost reasons, and others adopting a “wait-and-see” approach and expressing keen interest in practices from other agencies.

Figure 1.

Thirteen agencies selected for a review of tremie concrete practices.

There are two primary thermal requirements for concrete elements, each intended to prevent a different type of thermal damage that may compromise durability. The first requirement limits the maximum temperature within concrete elements. Such limits are intended to prevent delayed ettringite formation (DEF), a phenomenon that can lead to eventual expansion and cracking of concrete long after development of high temperatures during hydration. The second requirement limits the maximum temperature differential within concrete elements. Temperature differential limits are intended to prevent excessive thermal cracking that can result from the center and edges of a concrete element expanding and contracting at different rates. Both DEF and thermal cracking are addressed in this paper. For each, background information, agency practices, and recent research developments are summarized. Subsequently, a procedure for establishing thermal requirements to prevent DEF and excessive thermal cracking for drilled shaft concrete is presented. An example application of the procedure is also presented.

Thermal Requirements to Address Delayed Ettringite Formation (DEF)

The first of two primary durability concerns is DEF, which can produce long-term cracking in concrete elements that experience excessive temperatures during curing. Accordingly, DEF has typically been addressed by imposing allowable maximum temperature provisions.

Delayed Ettringite Formation (DEF)

Ettringite, the mineral term for calcium sulfoaluminate ( $3 CaO * A l_{2} O_{3} * 3 CaS O_{4} * 32 H_{2} O)$ , is typically found in Portland cement. Ettringite in its early form, commonly referred to as “primary ettringite,” is a needle-shaped crystal that benefits plastic concrete because of its ability to regulate early hydration reactions to prevent flash setting, increase early strength development, and reduce drying shrinkage ( 1 ). Ettringite forms when sulfate compounds, such as gypsum, react with calcium aluminate in the cement. Although ettringite formed at early ages is considered beneficial, high heats of hydration may lead to DEF, which is the “potentially deleterious reformation of ettringite in moist concrete, mortar, or paste after destruction of primary ettringite due to high temperatures” ( 2 ). Subsequent exposure of concrete to water enables the delayed formation of ettringite, potentially causing expansion and cracking of the concrete. The photo in Figure 2 is an example of cracking because of DEF.

Figure 2.

Cracking from distress caused by delayed ettringite formation (DEF) ( 3 ). (Reprinted, with permission, from “Evaluation and Mitigating Measures for Premature Concrete Distress in Texas Department of Transportation Concrete Elements,” Cement, Concrete, and Aggregates, 21(1), 1999, copyright ASTM International, 100 Barr Harbor Drive, West Conshohocken, PA 19428.)

In some cases, DEF can be considered a secondary mechanism of distress caused initially by alkali-silica reactivity (ASR). ASR is a reaction that occurs when alkalis ( $N a_{2} O$ ), found in the cement paste, react with siliceous aggregates to form an alkali-silica gel. This gel expands when water is absorbed, potentially cracking the surrounding paste ( 4 ). Use of aggregates with high alkali reactivity, such as limestone, can increase the risk of ASR, thereby increasing the risk of DEF.

It is important to emphasize that development of high temperatures in concrete elements does not necessarily mean that DEF will occur. Instead, temperature limits are imposed because limiting the maximum temperature for concrete will prevent DEF from occurring.

Agency Practices to Address DEF

Table 1 summarizes the maximum allowable temperature specified in mass concrete provisions for the four agencies that had applied mass concrete provisions to drilled shafts. Specifications for MNDOT and SCDOT impose a maximum temperature limit of 160°F, while FDOT allows a higher temperature of 180°F. Temperature limits for Caltrans depend on shaft diameter. For shafts less than 8 ft in diameter, no mass concrete provisions are applied. For shafts between 8 and 14 ft in diameter, maximum temperature is not specified, but is indirectly addressed by limitations imposed on the total cement content of the mix design. For shafts greater than 14 ft in diameter, allowable temperatures are established in project-specific thermal control plans.

Table 1.

Maximum Temperature Limits Specified by the Agencies Reviewed for Federal Highway Administration Study

Agency	Specification type	Maximum temperature
Caltrans	Standard specification	Depends on diameter
FDOT	Standard specification	180°F for shafts greater than 6 ft in diameter
MNDOT	Special provision	160°F
SCDOT	Special provision	160°F

Note: FDOT = Florida Department of Transport; MNDOT = Minnesota Department of Transport; SCDOT = South Carolina Department of Transport.

Prevention of DEF

Methods for mitigating risk of DEF include controlling temperatures during curing and increasing a concrete mix’s resistance to DEF. Methods for controlling temperatures that develop in concrete elements are numerous and can be grouped in three general strategies: (1) limit the heat of hydration, primarily through mix design, (2) reduce the temperature of the fresh concrete through temperature control methods, and (3) remove heat from the hydrating concrete with cooling techniques. Methods for increasing a concrete mix’s resistance to DEF are limited to mix design.

Unlike other DEF mitigation techniques, mix design provides an opportunity to prevent high temperatures altogether and create a mix with improved resistance to DEF, essentially providing opportunity for a “double-benefit.” Replacing some of the Portland cement with supplementary cementitious materials is recommended for reducing the heat of hydration without sacrificing workability while also increasing DEF resistance.

The American Concrete Institute (ACI) recommendations prescribe a maximum allowable concrete temperature of 158°F, which appears to be the primary source for most current agency temperature limits ( 5 ). However, in cases where peak temperature during curing is likely to exceed 158°F, ACI requires other preventative measures be taken to limit the potential for DEF as summarized in Table 2. These measures most commonly involve use of prescribed proportions of supplementary cementitious materials (SCMs), which explicitly recognizes the influence of mix design characteristics for preventing DEF. If the specific mix design characteristics as satisfied, the ACI recommendations allow the maximum allowable temperature to be increased from the 160°F criterion used by many agencies to 185°F. Peak temperatures greater than 185°F are not permissible.

Table 2.

American Concrete Institute (ACI) Recommended Measures for Reducing Potential for Delayed Ettringite Formation ( 5 ). (Authorized reprint from ACI 201.2R-16)

Maximum concrete temperature, $T$	Prevention required
$T \leq$ 158°F	No prevention required.
158°F $< T \leq$ 185°F	Use one of following to minimize risk of expansion: 1. Low-alkali Portland cement having moderate or high sulfate resistance (ASTM C150/C150M) 2. Portland cement with 1-day mortar strength ≤ 2850 psi (ASTM C109/C109M) 3. Portland cement in combination with the following SCMs a. ≥ 25% Class F fly ash (ASTM C618) b. ≥ 35% Class C fly ash (ASTM C618) c. ≥ 35% slag cement (ASTM C989/C989M) d. ≥ 5% silica fume (ASTM C1240) with ≥25% slag cement e. ≥ 5% silica fume (ASTM C1240) with ≥20% Class F fly ash f. ≥ 10% metakaolin (ASTM C618) 4. Blended hydraulic cement with SCM content listed in Item 3 (ASTM C595/C595M or ASTM C1157/C1157M).
$T >$ 185°F	Not permissible under any circumstances

Note: psi = pounds per square inch; SCM = supplementary cementitious material.

One or more of the ACI requirements for concrete with peak temperatures exceeding 158°F are commonly satisfied or nearly satisfied for many tremie concrete mix designs used for drilled shafts, including mixes classified as self-consolidating concrete. Table 2 suggests that the potential for DEF can be reduced without requiring substantial changes to existing concrete mix designs, and in turn allows a maximum temperature criterion to be set at 185°F. Adopting this approach would relax most current maximum allowable temperature criteria, often preventing the need to adopt other more disruptive and costly means to control temperatures.

Thermal Requirements to Address Thermal Cracking

The second of two primary durability concerns is thermal cracking: cracking that occurs in response to thermo-mechanical stresses resulting from the center and edges of a concrete element heating and cooling at different rates. Accordingly, thermal cracking has typically been addressed by imposing allowable temperature differential provisions.

Thermal Cracking

As concrete hydrates and heat accumulates, the surface of concrete elements generally cools at a faster rate than the interior as heat transfers to surrounding materials. As a result, the temperature of the concrete decreases with distance from the element center, which produces a temperature difference within the concrete member. Large temperature differences result in tensile stresses in the concrete, which can lead to thermal cracking.

Early age thermal cracking occurs because of the restraint concrete experiences while expanding or contracting during curing. Every concrete member has some form of restraint: without restraint, the concrete would develop no internal stress and would expand and contract freely without cracking ( 6 , 7 ). Restraint in concrete elements can be represented in design using a restraint factor, $R$ , defined as the ratio of the actual stress because of volume change for the element to the stress that would occur if the member was fully restrained. $R$ represents the degree of restraint or counterreaction, with $R = 0$ representing a concrete element free of stress and $R = 1$ representing a fully restrained element ( 8 ).

There are two categories of restraint, internal and external, both of which can contribute to cracking of early age concrete. Internal restraint of a concrete element occurs because of non-uniform volume change as a result of temperature differences within a concrete element. Figure 3 illustrates cracking because of internal restraint with time as curing progresses ( 7 ). The figure depicts a generic concrete element with steel reinforcing bars near the edges of the concrete. The left side of the diagram represents the time of concrete placement, with subsequent times being progressively illustrated to the right. During the heating phase, the magnitude of temperature differences increases as temperature in the interior of the element increases while the temperature of the element boundary remains at lower temperature. These temperatures cause the interior concrete to expand and “stretch” the element, while concrete near the boundary is contracting, or at least expanding less. The differential volume changes produced by the temperature differences cause tensile stresses and potentially cracking near the element boundary as shown in the left half of Figure 3. During subsequent cooling, concrete in the interior element begins to cool and contract while concrete on the element surface has reached a steady temperature. Contraction of concrete in the interior of the element causes concrete on the element surface to experience compression as tensile stresses build up in the element interior. These compressive stresses at the surface may cause previously developed cracks at the element surface to close, as shown to the right in Figure 3. Tension cracks may also form in the interior of the element at this time. Although Figure 3 depicts a generic concrete element, the figure and preceding discussion apply to drilled shafts.

Figure 3.

Conceptual illustration of crack development in early age concrete with time from internal restraint. (Adapted from Bamforth with permission from Construction Industry Research and Information Association [CIRIA] [7]).

External restraint results from external forces providing resistance to volume change, such as resistance imposed by the ground surrounding drilled shafts and other underground elements. During heating, external restraint provides resistance to cracking at the element surface as illustrated in Figure 4, and may even prevent cracking altogether. In contrast, during the cooling phase when the concrete is contracting and tensile stresses are developing in the interior of the element, external restraint contributes additional tensile stress causing increased thermal cracking ( 7 ). While drilled shafts and other underground elements are subject to external restraint along the element, internal restraint is believed to dominate thermal cracking for drilled shafts. For drilled shafts with permanent casing, thermal cracking is not a concern because the steel casing heats and cools with the surface of the concrete, while also confining the placement and restraining cracks from forming ( 9 ). Note that heat dissipation in drilled shafts is predominately radial, except near the ends of the shafts where longitudinal heat dissipation is also significant.

Figure 4.

Drilled shaft showing stress development during heating and cooling phases of concrete curing ( 1 ).

Agency Practices to Address Thermal Cracking

To limit thermal cracking, ACI 301 restricts the maximum allowable temperature difference between the center and surface of an element to be less than 35°F ( 10 ). Table 3 lists the maximum temperature differentials for the four agencies that had applied mass concrete provisions to drilled shafts. FDOT and SCDOT adopt the ACI 301 provision. MNDOT’s mass concrete special provision specifically states that maximum temperature differentials established for other mass concrete applications do not apply to drilled shafts. Caltrans does not limit temperature differentials for shafts less than 14 ft in diameter. For shafts with diameters greater than 14 ft, contractors must develop and submit a thermal control plan that includes maximum allowable temperature differentials to prevent cracking based on modeling results.

Table 3.

Maximum Temperature Differential Limits Specified by the Agencies Included in Federal Highway Administration Study

Agency	Specification type	Maximum temperature
Caltrans	Standard specification	8 ft < diameter ≤ 14 ft: not specified diameter > 14 ft: varies by project
FDOT	Standard specification	35°F
MNDOT	Special provision	Not specified
SCDOT	Special provision	35°F

Note: FDOT = Florida Department of Transport; MNDOT = Minnesota Department of Transport; SCDOT = South Carolina Department of Transport.

Recommended Procedure to Establish Allowable Temperature Differentials

Occurrence of thermal cracking is complicated because it depends on the magnitude of differential temperatures within a concrete element and the tensile strength of the concrete, both of which change with time, as well as the coefficient of thermal expansion of the concrete. Several approaches have been developed to predict thermal cracking in concrete. The approach described by Bamforth ( 7 ), developed for the UK-based Construction Industry Research and Information Association (CIRIA), was adopted for the current work, based on its practicality and relative simplicity. Other methods, including the method in ACI 207.2, were also considered, but are not the basis for the proposed methodology because they consider only temperature differentials associated with incipient cracking, not temperature differentials associated with limited cracking ( 6 ). Bamforth provides different equations for different forms of restraint; the equations for internal restraint are most relevant to drilled shafts and other below-ground elements and are summarized here ( 7 ). As described previously, however, deep foundation elements are also subject to restraint from the ground surrounding the element, which tends to further reduce cracking in concrete. The methods described here are therefore conservative.

Occurrence of thermal cracking is evaluated based on comparison of the “restrained strain,” $ε_{r},$ with the tensile strain capacity, $ε_{ctu}$ , where cracking is presumed to occur when $ε_{r}$ exceeds $ε_{ctu}$ . The restrained strain can be calculated as

ε_{r} = K \cdot Δ T \cdot α_{c} \cdot R

(1)

where $K$ = coefficient accounting for stress relaxation because of creep

$Δ T$ = temperature difference between center and outer surface

$α_{c}$ = coefficient of thermal expansion

$R$ = restraint factor

Restrained strain is mainly a function of the temperature difference, $Δ T$ , and the coefficient of thermal expansion, $α_{c}$ . The coefficient of thermal expansion for normal weight concrete varies between 3 to 8 $\times 10^{- 6}$ per °F depending primarily on the type of aggregate used as illustrated in Figure 5 ( 11 ). In the absence of more precise data, the coefficient of thermal expansion can be taken as $6.0 \times 10^{- 6}$ per °F ( 11 ). Both the coefficient, $K$ , and the restraint factor, $R$ , can be considered constants following conservative design recommendations provided in Bamforth ( 7 ). For cases of internal restraint that predominates cracking in drilled shafts, $R$ can be taken to equal 0.42 for all conditions. The coefficient $K$ can be taken to equal 0.65 for design purposes ( 7 ).

Figure 5.

Coefficients of thermal expansion for concrete with various aggregate types, based on ACI 207.2R-07 ( 6 ).

The tensile strain capacity, $ε_{ctu}$ , is the maximum strain the concrete can endure without continuous cracking. A lower-bound value that is appropriate for design can be calculated as in Bamforth ( 7 ):

ε_{ctu} = \frac{f_{t}}{E_{c}}

(2)

where $f_{t}$ = tensile strength

$E_{c}$ = modulus of elasticity

Both $f_{t}$ and $E_{c}$ can be directly measured; however, for design purposes, they are more commonly estimated from concrete compressive strength, $f_{c}^{'}$ ( 11 ).

If Equations 1 and 2 indicate thermal cracking will occur, acceptable durability of the concrete element can still be provided if the concrete element has sufficient longitudinal reinforcement to limit crack width to satisfy durability requirements. Crack width, $w_{k}$ , can be calculated as in Bamforth ( 7 )

w_{k} = ε_{cr} \cdot S_{r, \max} = ε_{cr} \cdot (3.4 c + 0.425 \frac{k_{1} k_{2} φ}{ρ_{p, eff}})

(3)

where $ε_{cr}$ = crack-inducing strain

$S_{r, \max}$ = maximum crack spacing

$c$ = concrete cover

$k_{1}$ = reinforcement bond coefficient (=0.8 for deformed bars)

$k_{2}$ = strain distribution coefficient (=1.0 for direct tension)

$φ$ = bar diameter

$ρ_{p, eff}$ = reinforcement ratio (= $A_{s} / A_{c, eff}$ )

$A_{c, eff}$ = effective area of concrete in tension around reinforcement

In Equation 3, the crack-inducing strain is computed as

ε_{cr} = ε_{r} - 0.5 ε_{ctu}

(4)

The tensile strain capacity, $ε_{ctu}$ , generally increases with time, because $f_{c}^{'}$ , $f_{t}$ , and $E_{c}$ increase with time as the concrete cures and gains strength.

For drilled shafts, the effective area of concrete in tension is computed as the product of the reinforcing cage circumference (from reinforcing cage diameter, $D_{cage}$ ) and effective depth of concrete in tension, $h_{c, eff}$ :

A_{c, eff} = π \cdot D_{cage} \cdot h_{c, eff}

(5)

The effective depth of concrete in tension is the lesser of the shaft radius, $r_{shaft}$ , or the quantity shown in Equation 6. Typically, the latter will be less than the shaft radius.

h_{c, eff} = \min (r_{shaft}, 2.5 \cdot (c + \frac{φ}{2}))

(6)

The methodology presented above can be used to establish maximum temperature differentials that account for appropriate concrete characteristics and are generally preferable to the simple constant values being employed by some agencies. One such expression can be developed considering maximum temperature differentials to prevent cracking altogether. Another expression can be developed relying on the benefit of reinforcing steel to limit cracking so that acceptable durability is maintained.

An expression for the maximum temperature differential to prevent cracking, $Δ T_{crack}$ , can be developed by equating Equations 1 and 2 and solving for the temperature differential, $Δ T$ . The resulting expression is

Δ T_{crack} = \frac{ε_{ctu}}{K \cdot α_{c} \cdot R} = \frac{3.7 \cdot ε_{ctu}}{α_{c}}

(7)

where $K = 0.65$ and $R = 0.42$ as described previously. Since $ε_{ctu}$ generally increases with time as concrete strengthens and cures, the temperature differential associated with cracking also increases with time.

An alternative, less restrictive expression for allowable temperature differentials can be developed if cracking is allowed, but limited to be less than some specific crack width to satisfy durability requirements. Such an expression can be established by setting Equation 3 equal to the limiting crack width, $w_{\lim}$ , and solving for the temperature differential, $Δ T_{limit}$ . The resulting expression in this case is

\begin{matrix} Δ T_{limit} & = \frac{(\frac{w_{\lim}}{3.4 c + 0.425 \frac{k_{1} k_{2} φ}{ρ_{p, eff}}}) + 0.5 \cdot ε_{ctu}}{K \cdot α_{c} \cdot R} \\ = \frac{(\frac{w_{\lim}}{3.4 c + 0.425 \frac{0.8 \cdot φ}{ρ_{p, eff}}}) + 0.5 \cdot ε_{ctu}}{0.27 \cdot α_{c}} \end{matrix}

(8)

where again $K$ is taken to be 0.65 and $R$ is taken to be 0.42, and the coefficients $k_{1}$ and $k_{2}$ are taken to be 0.8 and 1.0, respectively ( 7 ). The allowable temperature differential from Equation 8 also increases with time, and also increases with increasing limiting crack width, decreasing concrete cover ( $c$ ), and increasing reinforcement ratio ( $ρ_{p, eff}$ ). A value of $w_{\lim}$ equal to 0.3 mm is commonly adopted for concrete in contact with non-aggressive ground ( 7 , 12 ).

Recommended Procedure for Establishing and Satisfying Thermal Requirements for Drilled Shaft Concrete

The state of practice in relation to thermal requirements for drilled shafts generally consists of either neglecting the potential for thermal damage altogether or adopting provisions developed for above-ground concrete elements that are subject to considerably different thermal and mechanical boundary conditions. The procedure outlined below is a rational approach for considering thermal requirements for drilled shafts, adapting the concepts and techniques presented previously to drilled shafts.

Step 1: Define Input Parameters

The procedure involves three independent analyses: (a) predicting drilled shaft concrete temperatures, (b) establishing the allowable maximum temperature to prevent DEF, and (c) establishing the allowable temperature differential to limit thermal cracking to provide appropriate durability. Before performing the analyses in subsequent steps, it is necessary to gather information in relation to the input parameters required for the analyses. Required inputs are summarized in Table 4.

Table 4.

Summary of Required Inputs

Parameter	Analysis requirements
	Thermal model	Delayed ettringite formation (DEF) potential	Thermal cracking
Concrete mix design
Cement content and type of cement	X	X	X
Aggregate coefficient of thermal expansion	X	X	X
Concrete tensile strength	−	−	X
Concrete elastic modulus	−	−	X
Concrete placement temperature	X	−	−
Drilled shaft diameter	X	−	−
Drilled shaft reinforcement	−	−	X
Drilled shaft concrete cover distance	−	−	X
Soil or rock density	X	−	−
Soil or rock thermal conductivity	X	−	−
Soil or rock specific heat	X	−	−
Soil or rock temperature	X	−	−

Note: “X” indicates parameter is required for analysis; “–” indicates parameter is not required for analysis

Step 2: Predict Concrete Temperatures

To evaluate the potential for thermal damage, it is necessary to predict expected temperatures after concrete placement. Numerous techniques can be used to predict temperatures in drilled shafts. The methods range in complexity from simple chart solutions (e.g., the graphical method in ACI 207.2) through approximate numerical solutions (e.g., the Schmidt method, summarized in the 1996 version of ACI 207.1) to models based on numerical solution of the heat diffusion equation ( 6 , 13 ).

One finite difference solution technique is implemented in ConcreteWorks, a free computer program for analysis of mass concrete elements developed as part of research for the Texas Department of Transportation ( 14 ). ConcreteWorks includes modeling features for evaluating drilled shaft concrete elements, with the thermal properties for the surrounding geologic materials defined categorically by the type of material (e.g., sand, clay). El-tayash evaluated the use of ConcreteWorks for drilled shafts using thermal integrity profiling data and concluded ConcreteWorks is sufficiently accurate for estimating the thermal response of drilled shafts for different concrete mix parameters, element sizes, and ground conditions ( 1 ). Average differences between measured and predicted peak temperatures were approximately 12%. Predicted temperatures from ConcreteWorks modeling by El-tayash tended to exceed measured temperatures, in contrast with findings from previous studies focused on above-ground elements ( 15 , 16 ).

Regardless of the prediction technique, the required output from temperature prediction is the maximum temperature and the maximum temperature differential. Sensitivity analyses are prudent in many cases, especially because there is typically significant uncertainty for influential parameters such as soil and rock temperature, thermal conductivity, and specific heat. Uncertainty can be reduced by calibrating the thermal model using measured temperatures when measured temperatures are available.

Step 3: Address DEF Potential

The ACI guidance presented in Table 2 is recommended for addressing DEF potential. The peak temperature predicted from Step 2 should be used as $T$ in the table to determine if mix design revisions are necessary to reduce DEF potential. Incorporation of any mix design revisions requires repeating Step 2, since mix design revisions will alter temperature predictions.

Step 4: Establish Allowable Temperature Differential to Limit Concrete Cracking

Allowable temperature differential should be established using either Equation 7 if no cracking is permitted or Equation 8 if cracking is allowed but limited to some value ( $w_{\lim}$ ) to provide for appropriate durability. As noted previously, durability requirements can be satisfied while allowing for small crack widths. ACI 224R-01 suggests a limiting crack width of 0.3 mm for concrete exposed to soil, although smaller limiting crack widths may be considered in particularly aggressive ground conditions or based on consideration of anticipated service life ( 12 ). The approach of not permitting cracking (Equation 7) is likely overly restrictive for most applications.

The allowable temperature differential from Equations 7 or 8 depends on the tensile strain capacity, which in turn depends on concrete tensile strength and modulus of elasticity. Both parameters can be measured using concrete specimens. More commonly, the parameters are estimated from correlations with concrete compressive strength. Both AASHTO and the Eurocode include such correlations (11, 17). Because of uncertainty associated with the correlations, uncertainty in relation to compressive strength itself, and uncertainty in relation to how the parameters change with time, the concrete tensile strength and elastic modulus are significant sources of uncertainty for establishing the allowable temperature differential. The uncertainty can be reduced by using results of measurements of compressive strength rather than the design value or, more effectively, by actually measuring concrete tensile strength, elastic modulus, or both. Developing strength gain curves from compressive strength tests at different points in time (e.g., 3-day, 7-day, 28-day) can reduce uncertainty in relation to how the parameters change with time.

The allowable temperature differential from Equations 7 and 8 is also highly influenced by the coefficient of thermal expansion. Because of the large range of potential values of the coefficient of thermal expansion, laboratory measurement for specific concrete mixes can significantly reduce uncertainty in predictions of allowable temperature differentials.

Step 5: Compare Predicted and Allowable Temperatures

The predicted maximum temperature differential from Step 2 should be compared with the allowable value from Step 4. If the predicted value exceeds the allowable value, mitigation techniques via Step 6 are necessary. The comparison considers only temperature differential and not maximum temperature, since DEF is presumed to have been addressed via Step 3. In addition, if predicted temperatures from Step 2 are relatively close to the threshold values from Steps 3 and 4—say maximum temperatures greater than 150°F or temperature differential values within 75% of the allowable value—it is recommended that temperatures within the drilled shaft be measured via Step 7.

Step 6: Mitigate Excessive Temperature Differentials (if Necessary)

Recommended mitigation techniques for drilled shaft concrete include mix design modifications and techniques to limit concrete placement temperature. Since concrete mix parameters strongly influence the temperatures that develop as well as the resistance to DEF and thermal cracking, modifications to the concrete mix can be quite effective in preventing DEF and limiting concrete cracking. Potential mix design modifications to reduce temperatures, increase resistance to DEF and thermal cracking, or both, are listed below:

Replace some portion of Portland cement with fly ash or slag cement. For reasons explained in ACI 201.2R-16 and per Table 2, use of fly ash and slag cement increases resistance to DEF and therefore increases the allowable maximum temperature ( 5 ). Use of fly ash and slag cement also generally reduces the heat of hydration, which reduces concrete temperature and is therefore beneficial for preventing both DEF and thermal cracking. Fly ash and slag cement are also generally associated with increased strength and reduced permeability, both of which positively influence durability of constructed concrete elements.

Use aggregate with lower coefficient of thermal expansion. The aggregate coefficient of thermal expansion strongly influences the potential for thermal cracking. Using aggregate sources with relatively small coefficient of thermal expansion, for example, limestone, can greatly reduce thermal cracking and therefore increase the allowable temperature differential.

Reduce total cement content. Reducing cement content lowers temperatures and therefore reduces the potential for both DEF and thermal cracking. However, reducing cement content will also reduce strength and potentially stability and workability, and should therefore be used with caution.

As noted throughout the list above, designers must consider the implications of mix design modifications, particularly with respect to concrete workability and strength. In addition, agency specifications may present administrative challenges to mix design modifications.

Reducing concrete placement temperature is among the most effective means for reducing internal temperatures within the concrete: Mehta and Monteiro report that every 10°F reduction in placement temperature can reduce the maximum temperature that will develop within the concrete by an average of 6°F ( 18 ). There are various techniques that can be used to limit placement temperature:

Batching and placing concrete at night when mix components are generally coolest

Flushing aggregate with cold water

Using chilled mix water, replacing some of the mix water with crushed ice, or both

Using liquid nitrogen to cool the final mix

In above-ground concrete, post-cooling techniques that entail flushing cold fluid through pipes cast into the concrete member are sometimes used. Post-cooling techniques should generally be avoided for drilled shaft concrete unless other options are impractical. Placement of concrete for large-diameter drilled shafts is sufficiently challenging without additional congestion because of cooling pipes. Moreover, the congestion because of cooling pipes, which must be in the center of the shaft to effectively limit peak temperatures and temperature differentials, is more detrimental to tremie concrete placement than the congestion typically encountered at the reinforcing cage.

Step 7: Measure Temperatures (if Necessary)

If the evaluation of Step 5 results in a recommendation for monitoring temperatures, instrumentation should be included within the drilled shaft concrete. Temperature measurements should be recorded at the center and near the edge of the drilled shaft at a depth of at least one diameter below the ground surface. The measured temperatures should be compared with allowable values from Steps 3 and 4: if the measured temperatures exceed the allowable values, mitigation techniques from Step 5 should be implemented. Calibrating (or re-calibrating) the thermal model from Step 2 using measured temperatures will likely help to evaluate mitigation techniques, and may also prove useful for future projects implementing the same concrete mix or installed in similar ground conditions.

Example Application of Recommended Procedure

The recommended procedure was applied to 10 ft diameter drilled shafts for a transportation project in the Pacific Northwest. For 10 of the shafts, temperature measurements were collected near the center and edge for the first 10 days following concrete placement. Each step of the procedure is summarized below.

Step 1: Define Input Parameters

Input parameters for the project are summarized in Table 5. The concrete mix includes 35% slag cement in anticipation of high temperatures that would trigger DEF mitigation provisions. The high temperatures were anticipated for the large-diameter shafts based on local experience with typical mix designs in the area, which include only Portland cement. The contractor reported significant problems with concrete pumpability during installation of a technique shaft, requiring re-proportioning the mix to include more sand and less coarse aggregate. The pumpability issues were attributed to the slag cement. Use of fly ash rather than slag cement to limit temperatures and resist DEF would likely have prevented workability issues, but fly ash was not readily available near the project site.

Table 5.

Summary of Input Parameters for Example Project

Parameter	Value	Notes
Concrete mix design	390 lb Type IL cement 210 lb slag cement 242 lb water 1,615 lb sand (siliceous river sand) 1,615 lb 3/8 in. coarse aggregate (siliceous river gravel) High-range water reducer Viscosity modifying admixture Hydration control 2% Entrained air	Coefficient of thermal expansion unknown; conservatively use 8µε/°F w/c = 0.40 Cement content is 35% slag cement
Concrete compressive strength	$f'_{c} = 2, 500 \ln (t) + 1, 500$	$f'_{c}$ in psi, $t$ in days Regression based on 2-day, 7-day, and 28-day measurements
Concrete tensile strength	Correlation with $f'_{c}$	AASHTO LRFD Bridge Specs ( 11 )
Concrete elastic modulus		Eurocode BS EN 1992-1-1 ( 17 )
Concrete placement temperature	70°F	Actual
	85°F	Worst-case
Drilled shaft diameter	10 ft	Design diameter
Drilled shaft reinforcement	27 bundles of 2 #11 bars	Lightest reinforcement among various cage designs
Drilled shaft concrete cover distance	8 in.	Largest value among various cage designs
Allowable crack width	0.3 mm	ACI 224 ( 12 )
Soil density	2320 kg/m³	See Step 2; model inputs calibrated to match measured temperatures
Soil thermal conductivity	2.15 W/(m·K)	See Step 2; model inputs calibrated to match measured temperatures
Soil specific heat	810 J/(kg·K)	See Step 2; model inputs calibrated to match measured temperatures
Soil or rock temperature	55°F	Except near the surface, soil temperature is controlled by groundwater temperature, which is typically about 55°F near the project site

Note: psi = pounds per square inch; w/c = water-to-cementitious materials ratio.

Step 2: Predict Concrete Temperatures

ConcreteWorks was used to model the thermal response of the drilled shafts. Model inputs were consistent with Table 5. ConcreteWorks inputs for soil density, thermal conductivity, and specific heat cannot be entered directly; rather, they are defined based on a dropdown list of geologic materials. The properties listed in Table 5 are for limestone, which is inconsistent with sandy clay geology at the project site. Although limestone is inconsistent with project geology, limestone was selected during model calibration using measured temperatures.

Results of the analyses are shown in Figure 6, which shows temperature at the center of the shaft versus time, and in Figure 7, which shows temperature differential versus time.

Figure 6.

Temperature at center of shaft versus time from thermal model and from temperature measurements.

Figure 7.

Temperature differential between center and edge of shaft versus time from thermal model and from temperature measurements.

The results indicate strong agreement between ConcreteWorks predictions for the 70°F placement temperature model and observed temperatures. At a maximum of about 150°F, the predicted center temperature is essentially equal to the average of observed temperatures for the observation period of approximately 225 h. At a maximum of about 35°F, the predicted temperature differential also agrees closely with the average value of observations, although the model predicts differentials dissipating somewhat more quickly than was observed.

It is also important to consider the variability of observed results, as well as ConcreteWorks predictions for analyses using the 85°F placement temperature. For the temperature at the center of the shaft, the average peak value is approximately 150°F, but the data vary from about 135°F to 160°F. The ConcreteWorks model for 85°F placement temperature predicts a maximum temperature of 171°F. For the temperature differential, the average value is about 35°F, but the data range from 24°F to 46°F. The ConcreteWorks model for 85°F placement temperature predicts a maximum temperature differential of 40°F.

Step 3: Address DEF Potential

As described in Step 1, the total cementitious content of the mix design includes 35% slag cement to satisfy ACI requirements summarized in Table 2. As presented in Figure 6, the thermal model and temperature measurements indicate the ACI requirements were satisfied.

Step 4: Establish Allowable Temperature Differential

The inputs defined in Table 5 were used to calculate the allowable temperature differential per Equation 8. The value of coefficient of thermal expansion of 8µε/°F is likely conservative, but was used because of limited information in relation to local aggregate. The resulting allowable temperature differential was computed to be 86°F.

Step 5: Compare Predicted and Allowable Temperatures

The predicted temperatures and temperature differentials from Step 2 satisfy requirements for DEF from ACI (Table 2) and are less than the allowable temperature differential based on recommendations provided in this paper, so no mitigation measures are required. The predicted maximum temperature is 150°F, which is greater than 75% of the allowable maximum temperature, so temperature measurements are recommended.

Step 6: Mitigate Excessive Temperature Differentials

Although the results presented in previous steps indicate mitigation measures are not required, mitigation measures were preemptively included based on local experience as described in Step 1. Measures to limit placement temperature were also proactively adopted, including using chilled mix water.

Step 7: Measure Temperatures

Based on the Step 5 assessment, temperatures were monitored. Results of the monitoring are included in Figures 6 and 7. The results indicate the drilled shaft concrete thermal response was consistent with project requirements. The thermal response should not produce any durability concerns.

Conclusions

Consideration of experience with thermal performance of above-ground concrete and the magnitude of temperatures commonly observed in drilled shaft concrete, particularly large-diameter drilled shafts, suggests it is prudent to evaluate thermal requirements when designing drilled shafts for durability. However, consideration of currently adopted criteria and thermal cracking, as well as unique characteristics of drilled shafts, suggests the provisions that some agencies have imposed on drilled shafts are overly restrictive.

In particular, the commonly adopted maximum temperature differential of 35°F is likely much more stringent than necessary to provide satisfactory durability for most drilled shafts. In reality, large-diameter shafts are typically heavily reinforced, confined by soil, rock, or both, and can withstand small cracks while still providing satisfactory durability.

The procedure recommended in this paper establishes a rational methodology for addressing DEF and thermal cracking based on explicit consideration of the thermo-mechanical response of concrete for predicted temperatures. The procedures allow for explicit account of project-specific characteristics, including ground conditions, concrete mix design characteristics, drilled shaft geometry, and the quantity of steel reinforcement. The methodology was developed from guidance established by ACI and CIRIA and provides a rational means for designing drilled shafts for durability without imposing unnecessary constraints that may exacerbate challenges with effective construction of drilled shafts.

Footnotes

Acknowledgements

Guidance from Silas Nichols and Justice Maswoswe of FHWA was especially helpful. The authors also appreciate thoughtful administration of the research from WSP, and from Brian Zelenko in particular.

Author Contributions

The authors confirm contribution to the paper as follows: study conception and design: A. Boeckmann, Z. El-tayash, J. Loehr; data collection: A. Boeckmann, Z. El-tayash, J. Loehr; analysis and interpretation of results: A. Boeckmann, Z. El-tayash, J. Loehr; draft manuscript preparation: A. Boeckmann, Z. El-tayash, J. Loehr. All authors reviewed the results and approved the final version of the manuscript.

Declaration of Conflicting Interests

The author(s) declared no potential conflicts of interest with respect to the research, authorship, and/or publication of this article.

Funding

The author(s) disclosed receipt of the following financial support for the research, authorship, and/or publication of this article: The authors gratefully acknowledge the Federal Highway Administration for funding the work described in this paper, which was performed under contract number DTFH6114R00006.

References

El-tayash

. Evaluation of Temperature Predictions and Thermal Criteria for Design of Drilled Shaft Foundations. Thesis. University of Missouri, 2021.

Portland Cement Association. Ettringite Formation and the Performance of Concrete. IS417. Portland Cement Association, Skokie, IL, 2001, pp. 1–16.

Lawrence

B. L.

Moody

E. D.

Guillemette

R. N.

Carrasquillo

R. L.

Evaluation and Mitigating Measures for Premature Concrete Distress in Texas Department of Transportation Concrete Elements. Cement, Concrete, and Aggregates, Vol. 21, No. 1, 1999, pp. 73–81.

Bauer

Cornell

Figurski

Ley

Miralles

Alkali-Silica Reaction and Delayed Ettringite Formation in Concrete: A Literature Review. Report. University of Texas, Austin, TX, 2006, p. 61.

ACI 201.2R-16. Guide to Durable Concrete. ACI 201.2R-16. American Concrete Institute, Farmington Hills, MI, 2016.

ACI 207.2R-07. Thermal and Volume Change Effects on Cracking of Mass Concrete. ACI 207.2R-07. American Concrete Institute, Farmington Hills, MI, 2007.

Bamforth

P. B.

Control of Cracking Caused by Restrained Deformation in Concrete. C766. CIRIA, London. http://www.ciria.org.

Al-Gburi

Restraint Effects in Early Age Concrete Structures. Doctoral thesis. Luleå tekniska universitet, Sweden, 2015.

Gajda

Drilled Shafts as Mass Concrete? Concrete Bridge Views, 2016.

10.

ACI 301R-10. Specifications for Structural Concrete. ACI 301R-10. American Concrete Institute, Farmington Hills, MI, 2010.

11.

AASHTO. AASHTO LRFD Bridge Design Specification: Customary U.S. Units, 7th ed. American Association of State Highway and Transportation Officials, Washington, D.C., 2014.

12.

ACI 224R-01. Control of Cracking in Concrete Structures. ACI 224R-01. American Concrete Institute, Farmington Hills, MI, 2001.

13.

ACI 207.1R-96. Mass Concrete. ACI207.1R-96. American Concrete Institute, Farmington Hills, MI, 1996.

14.

Riding

Schindler

Pesek

Drimalas

Folliard

ConcreteWorks V3 Training/User Manual. Report 0-6332-P1. Center for Transportation Research, University of Texas, Austin, TX, 2017.

15.

Shaw

J. J.

Jahren

C. T.

Wang

Iowa Mass Concrete for Bridge Foundations Study – Phase II. Final Report. Iowa State University Institute for Transportation, Ames, IA, 2014.

16.

Gross

Development of a Mass Concrete Specification for Use in ALDOT Bridge Construction. Master’s thesis. Auburn University, AL, 2017.

17.

EN1992-1-1. Eurocode 2: Design of Concrete Structures – Part 1-1: General Rules and Rules for Buildings. The European Union, 2004.

18.

Mehta

P. K.

Monteiro

P. J. M.

Concrete: Microstructure, Properties, and Materials, 4th ed. McGraw-Hill Education, 2014.