109-05-aci materials journal sept.-oct. 2012 complete

ACIMATERIALS JOURNAL

VOL. 109, NO. 5SEPTEMBER-OCTOBER 2012

499 Creep Testing of Epoxy-Bonded Reinforcing Bar Couplers/G. Brungraber

503 Effect of Na2SiO

3/NaOH Ratios and NaOH Molarities on Compressive Strength

of Fly-Ash-Based Geopolymer/A. M. Mustafa Al Bakri, H. Kamarudin, M. Bnhussain, A. R. Rafiza, and Y. Zarina

509 Effect of Using Mortar Interface and Overlays on Masonry Behavior by Using Taguchi Method/M. Farouk Ghazy

517 Experimental Study on Dynamic Axial Tensile Mechanical Properties of Concrete and Its Components/S. Wu, Y. Wang, D. Shen, and J. Zhou

529 Potential Recycling of Bottom and Fly Ashes in Acoustic Mortars and Concretes/C. Leiva, L. F. Vilches, C. Arenas, S. Delgado, and C. Fernández-Pereira

537 Early-Age Creep of Mass Concrete: Effects of Chemical and Mineral Admixtures/S. Botassi dos Santos, L. C. Pinto da Silva Filho, and J. L. Calmon

545 Proposed Flexural Test Method and Associated Inverse Analysis for Ultra-High-Performance Fiber-Reinforced Concrete/F. Baby, B. Graybeal, P. Marchand, and F. Toutlemonde

557 A First-Cut Field Method to Evaluate Limestone Aggregate Durability/ J. R. Emry, R. H. Goldstein, and E. K. Franseen

565 Investigation of Properties of Engineered Cementitious Composites Incorporating High Volumes of Fly Ash and Metakaolin/E. Özbay, O. Karahan, M. Lachemi, K. M. A. Hossain, and C. Duran Atis

573 Fatigue Analysis of Plain and Fiber-Reinforced Self-Consolidating Concrete/ S. Goel, S. P. Singh, and P. Singh

A JOURNAL OF THE AMERICAN CONCRETE INSTITUTE

ACI Materials Journal/September-October 2012 497

Discussion is welcomed for all materials published in this issue and will appear in the July-August 2013 issue if received by April 1, 2013. Discussion of material received after specified dates will be considered individually for publication or private response.ACI Standards published in ACI Journals for public comment have discussion due dates printed withthe Standard.Annual index published online at www.concrete.org/pubs/journals/mjhome.asp.

ACI Materials JournalCopyright © 2012 American Concrete Institute. Printed in the United States of America.

The ACI Materials Journal (ISSN 0889-325x) is published bimonthly by the American Concrete Institute. Publication office: 38800 Country Club Drive, Farmington Hills, MI 48331. Periodicals postage paid at Farmington, MI, and at addi-tional mailing offices. Subscription rates: $161 per year (U.S. and possessions), $170 (elsewhere), payable in advance. POSTMASTER: Send address changes to: ACI Materials Journal, P. O. Box 9094, Farmington Hills, MI 48333-9094.

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CONTENTSBoard of Direction

PresidentJames K. Wight

Vice PresidentsAnne M. EllisWilliam E. Rushing Jr.

DirectorsNeal S. AndersonKhaled AwadRoger J. BeckerJeffrey W. ColemanRobert J. FroschJames R. HarrisCecil L. JonesSteven H. KosmatkaDavid A. LangeDenis MitchellJack P. MoehleDavid H. Sanders

Past President Board MembersKenneth C. HoverFlorian G. BarthLuis E. García

Executive Vice PresidentRon Burg

Technical Activities CommitteeDavid A. Lange, ChairDaniel W. Falconer, SecretarySergio M. AlcocerJoAnn P. BrowningChiara F. FerrarisCatherine E. FrenchTrey HamiltonRonald JanowiakKevin A. MacDonaldAntonio NanniJan OlekMichael SprinkelPericles C. StivarosEldon G. Tipping

StaffExecutive Vice PresidentRon Burg

EngineeringManaging DirectorDaniel W. Falconer

Managing EditorKhaled Nahlawi

Staff EngineersMatthew R. SenecalGregory Zeisler

Publishing ServicesManagerBarry M. Bergin

EditorsCarl R. BischofKaren CzedikKelli R. SlaydenDenise E. Wolber

Publishing AssistantAshley Poirier

ACI MAterIAls JournAl

septeMber-oCtober 2012, V. 109, no. 5a journal of the american concrete institutean international technical society

499 Creep Testing of Epoxy-Bonded Reinforcing Bar Couplers by Griffin Brungraber

503 Effect of Na2SiO3/NaOH Ratios and NaOH Molarities on Compressive Strength of Fly-Ash-Based Geopolymer by A. M. Mustafa Al Bakri, H. Kamarudin, M. Bnhussain, A. R. Rafiza, and Y. Zarina

509 Effect of Using Mortar Interface and Overlays on Masonry Behavior by Using Taguchi Method by Mariam Farouk Ghazy

517 Experimental Study on Dynamic Axial Tensile Mechanical Properties of Concrete and Its Components by Shengxing Wu, Yao Wang, Dejian Shen, and Jikai Zhou

529 Potential Recycling of Bottom and Fly Ashes in Acoustic Mortars and Concretes by Carlos Leiva, Luis F. Vilches, Celia Arenas, Silvia Delgado, and Constantino Fernández-Pereira

537 Early-Age Creep of Mass Concrete: Effects of Chemical and Mineral Admixtures by Sergio Botassi dos Santos, Luiz Carlos Pinto da Silva Filho, and João Luiz Calmon

545 Proposed Flexural Test Method and Associated Inverse Analysis for Ultra-High-Performance Fiber-Reinforced Concrete by Florent Baby, Benjamin Graybeal, Pierre Marchand, and Francois Toutlemonde

557 A First-Cut Field Method to Evaluate Limestone Aggregate Durability by Julienne Ruth Emry, Robert H. Goldstein, and Evan K. Franseen

565 Investigation of Properties of Engineered Cementitious Composites Incorporating High Volumes of Fly Ash and Metakaolin by E. Özbay, O. Karahan, M. Lachemi, K. M. A. Hossain, and C. Duran Atis

573 Fatigue Analysis of Plain and Fiber-Reinforced Self-Consolidating Concrete by S. Goel, S. P. Singh, and P. Singh

583 In ACI Structural Journal

498 ACI Materials Journal/September-October 2012

Permission is granted by the American Concrete Institute for libraries and other users registered with the Copyright Clearance Center (CCC) to photocopy any article contained herein for a fee of $3.00 per copy of the article. Payments should be sent directly to the Copyright Clearance Center, 21 Congress Street, Salem, MA 01970. ISSN 0889-3241/98 $3.00. Copying done for other than personal or internal reference use without the express written permission of the American Concrete Institute is prohibited. Requests for special permission or bulk copying should be addressed to the Managing Editor, ACI Materials Journal, American Concrete Institute.

The Institute is not responsible for statements or opinions expressed in its publications. Institute publications are not able to, nor intend to, supplant individual training, responsibility, or judgment of the user, or the supplier, of the information presented.

Papers appearing in the ACI Materials Journal are reviewed according to the Institute’s Publication Policy by individuals expert in the subject area of the papers.

MEETINGSContributions to ACI Materials Journal

The ACI Materials Journal is an open forum on concrete technology and papers related to this field are always welcome. All material submitted for possible publi-cation must meet the requirements of the “American Concrete Institute Publi-cation Policy” and “Author Guidelines and Submission Procedures.” Prospective authors should request a copy of the Policy and Guidelines from ACI or visit ACI’s Web site at www.concrete.org prior to submitting contributions.

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The Institute reserves the right to return, without review, contributions not meeting the requirements of the Publication Policy.

All materials conforming to the Policy requirements will be reviewed for editorial quality and technical content, and every effort will be made to put all acceptable papers into the information channel. However, potentially good papers may be returned to authors when it is not possible to publish them in a reasonable time.

Discussion

All technical material appearing in the ACI Materials Journal may be discussed. If the deadline indicated on the contents page is observed, discussion can appear in the designated issue. Discussion should be complete and ready for publication, including finished, reproducible illustra-tions. Discussion must be confined to the scope of the paper and meet the ACI Publi-cation Policy.

Follow the style of the current issue. Be brief—1800 words of double spaced, typewritten copy, including illustrations and tables, is maximum. Count illustrations and tables as 300 words each and submit them on individual sheets. As an approximation, 1 page of text is about 300 words. Submit one original typescript on 8-1/2 x 11 plain white paper, use 1 in. margins, and include two good quality copies of the entire discussion. References should be complete. Do not repeat references cited in original paper; cite them by original number. Closures responding to a single discussion should not exceed 1800-word equivalents in length, and to multiple discussions, approximately one half of the combined lengths of all discussions. Closures are published together with the discussions.

Discuss the paper, not some new or outside work on the same subject. Use references wherever possible instead of repeating available information.

Discussion offered for publication should offer some benefit to the general reader. Discussion which does not meet this requirement will be returned or referred to the author for private reply.

Send manuscripts to:http://mc.manuscriptcentral.com/aci

Send discussions to:[email protected]

2012

SEPTEMBER

19-21—4th International Conference on Accelerated Pavement Testing, Davis, CA, www.ucprc.ucdavis.edu/APT2012

19-21—18th IABSE Congress,s Seoul, South Korea, www.iabse.org/Seoul2012

19-21—8th RILEM International Symposium on Fibre Reinforced Concrete, Guimarães, Portugal, www.befib2012.civil.uminho.pt

20-23—ASCC Annual Conference, Chicago, IL, www.ascconline.org

21-23—4th International Conference on Problematic Soils, Wuhan, China, www.cipremier.com/page.php?487

24-28—15th World Conference on Earthquake Engineering, Lisbon, Portugal, www.15wcee.org

SEPTEMBER/OCTOBER

29-2—PCI Annual Convention & Exhibition and National Bridge Conference, Nashville, TN, www.pci.org

OCTOBER

2-4—Tilt-Up Concrete Association Annual Convention, Amelia Island, FL, www.tilt-up.org

3-6—NCPA 47th Annual Convention, New Orleans, LA, www.precast.org/convention

11-14—International Concrete Polishing & Staining Conference, Duluth, GA, www.icpsc365.com/icpsc2011

22-23—Building Envelope Technology Symposium, Phoenix, AZ, www.rci-online.org/symposium.html

28-31—Tenth International Conference on Superplasticizers and Other Chemical Admixtures in Concrete, Prague, Czech Republic, www.intconference.org

OCTOBER/NOVEMBER

31-2—Twelfth International Conference on Recent Advances in Concrete Technology and Sustainability Issues, Prague, Czech Republic, www.intconference.org

NOVEMBER

3-8—International Pool | Spa | Patio Expo, New Orleans, LA, www.poolspapatio.com

4-7—First International Conference for PhD Students in Civil Engineering, Cluj-Napoca, Romania, http://sens-group.ro/ce2012

11-14—Bridges Middle East, Doha, Qatar, www.bridgesme.com

14-16—Greenbuild 2012, San Francisco, CA, www.greenbuildexpo.org

UPCOMING ACI CONVENTIONSThe following is a list of scheduled ACI conventions:

2012—October 21-25, Sheraton Centre, Toronto, ON, Canada2013—April 14-18, Hilton & Minneapolis Convention Center, Minneapolis, MN2013—October 20-24, Hyatt Regency & Phoenix Convention Center, Phoenix, AZ2014—March 23-27, Grand Sierra Resort, Reno, NV

For additional information, contact:Event Services, ACI38800 Country Club DriveFarmington Hills, MI 48331Telephone: (248) 848-3795e-mail: [email protected]


Title no. 109-M47

ACI MATERIALS JOURNAL TECHNICAL PAPER

ACI Materials Journal, V. 109, No. 5, September-October 2012.MS No. M-2010-154.R1 received May 25, 2011, and reviewed under Institute

publication policies. Copyright © 2012, American Concrete Institute. All rights reserved, including the making of copies unless permission is obtained from the copyright proprietors. Pertinent discussion including author’s closure, if any, will be published in the July-August 2013 ACI Materials Journal if the discussion is received by April 1, 2013.

Creep Testing of Epoxy-Bonded Reinforcing Bar Couplersby Griffin Brungraber

test exists for reinforcing bar couplers; however, a creep test for adhesive concrete anchors is applicable due to their phenomenological similarity to epoxy-bonded reinforcing bar couplers.

MaterialsSystem C is a two-component structural adhesive. It is a

solvent-free, nonshrink, nonsag anchoring compound. The mixing ratio of the system is 1:1, resin:hardener. The resin and hardener are dispensed from a dual-cartridge system and simultaneously mixed in a static mixing nozzle. The system meets ASTM C881/C881M-02; Types I, II, IV, and V; Grade 3; Classes A, B, and C.1

The first component of System C is composed primarily of a bisphenol A-epichlorohydrin diepoxy resin and neopentyl glycol diglycidyl ether mixture; together these compounds act as the epoxy resin. The second component is composed primarily of n-aminoethyl piperazine and a nonylphenol mixture; together these compounds act as an amine adduct.

System F is a two-component, 100% solids, structural epoxy. When mixed, the resin and hardener combine into a smooth, nonabrasive paste adhesive. The mixing ratio of the system is 1:1, resin:hardener. The resin and hardener are dispensed from a dual-cartridge system and simultane-ously mixed in a static mixing nozzle. The system meets ASTM C881/C881M; Types I, II, IV, and V; Grade 3; Classes A, B, and C, except gel times.1 The gel time for the system varies from 10 to 30 minutes at temperatures from 4 to 32°C (39.2 to 89.6°F), respectively.

The first component of System F is composed primarily of bisphenol A-epichlorohydrin diepoxy resin; the component also contains small portions of methyl toluenesulfonate. The second component is composed primarily of piperazinyl-ethylamine and nonylphenol; the component also contains a proprietary mixture of fillers.

SpecimensNumber 5 (No. 16 metric) epoxy-bonded reinforcing bar

couplers were assembled with two different epoxy systems. Two specimens were assembled with each epoxy system for a total of four test specimens. Specimen assembly, curing, and testing were performed at a temperature of 23°C (73°F) and a relative humidity of 50%. The test setup is shown sche-matically in Fig. 1 and in a photograph in Fig. 2.

Experimental rationaleCreep of adhesive concrete anchors is typically evalu-

ated using ASTM E1512-01(2007)2; therefore, the creep test

A 40-day creep test was performed on epoxy-bonded reinforcing bar coupler specimens to assess their resistance to sustained tensile load. The reinforcing bar couplers were assembled using two different epoxy products. One was the product specified by the reinforcing bar coupler manufacturer; the other was a similar product that had already shown susceptibility to long-term failure in its intended application. The difference in the creep perfor-mance of the epoxy-bonded reinforcing bar couplers assembled with different epoxy products was significant—the manufacturer-specified epoxy product produced acceptable creep performance and the other epoxy product was shown to creep extensively.

Keywords: adhesive; anchorage; coupler; creep; epoxy; reinforcing bar; splice.

INTRODUCTIONReinforcing bar couplers are used to transfer load between

reinforcing bars in concrete structures in situations for which a typical lap splice would be inconvenient, inefficient, or inappropriate. Epoxy-bonded reinforcing bar couplers are one type of reinforcing bar coupler that are commercially available and unique in that they transfer load via an epoxy-filled sleeve. Although not all adhesives are epoxies, only epoxy has been used in adhesively bonded reinforcing bar couplers; therefore, the term “epoxy-bonded reinforcing bar coupler” is used to describe a class of construction products, although the mechanisms it refers to would be applicable to any adhesively bonded reinforcing bar coupler.

Adhesives, although commonly used in other branches of engineering, are relatively new to civil engineering (compared to materials such as steel and concrete). Conse-quently, their long-term degradation mechanisms are not as well-understood by the civil engineers responsible for their design as part of a reinforced concrete structure.

Epoxies, like many adhesives, are typically vulnerable to moisture and elevated temperature—two environmental conditions that are commonly found inside reinforced concrete structures, such as bridge decks.

RESEARCH SIGNIFICANCEEpoxy-bonded reinforcing bar couplers are commercially

available and have been installed in numerous reinforced concrete structures. Due to their reliance on an adhesive to transfer load—unique among reinforcing bar couplers—they may be susceptible to the same type of creep fail-ures that have been seen in epoxy anchorage to concrete. This research demonstrates the difference in performance between two commercially available epoxy systems. The results show that the creep performance of epoxy-bonded reinforcing bar couplers varies widely, depending on the epoxy system used.

EXPERIMENTAL INVESTIGATIONReinforcing bar coupler specimens were assembled using

two different types of epoxy. The only variable in their assembly was the epoxy system used. No standardized creep


creep of reinforcing bar couplers, the results of this test were compared to the results of a test on cementitiously grouted reinforcing bar couplers by Jansson.3

ANALYTICAL PROCEDUREThe results of the test were used to evaluate the perfor-

mance of epoxy-bonded reinforcing bar couplers. The analytical procedure for extrapolation of the creep data was taken from ASTM E1512. Data were recorded for 42 days and the results for each system were averaged between the specimens. Figure 3 shows the displacement time test data for both epoxy systems on a logarithmic time scale and fits the data to

( )lny c x b= ⋅ + (1)

where y is displacement; c is a curve-fitting constant; x is time; and b is a curve-fitting constant.

The curves developed from the data are extrapolated to a time of 600 days, as per ASTM E1512. Figure 4 shows the displacement time test data and associated curve fits extrap-olated to 600 days on a linear time scale.

ACI member Griffin Brungraber is an Assistant Bridge Engineer at T.Y. Lin Interna-tional. He received his BS from Bucknell University, Lewisburg, PA, and his MS and PhD from the University of California, San Diego, La Jolla, CA. He is a member of ACI Committee 355, Anchorage to Concrete. His research interests include the long-term performance of adhesives for concrete construction, infrastructure durability, and the seismic performance of bridges.

setup for the epoxy-bonded reinforcing bar couplers adapted ASTM E1512 wherever possible. The testing procedure for ASTM E1512 is to apply constant load to the test specimens and measure displacement. The load was initially applied over the course of 1 minute and then maintained for 42 days. The 42-day length of the test, the extrapolation of data to a time of 600 days, and the stress level of the test—40% of reinforcing bar ultimate strength—were all adapted from ASTM E1512. Although no specific standard exists for the

Fig. 1—Schematic figure of reinforcing bar coupler test setup.

Fig. 2—Photograph of reinforcing bar coupler test setup.

Fig. 3—Displacement time creep test data and fitted curves.

Fig. 4—Displacement time creep test data and fitted curves extrapolated to 600 days.


EXPERIMENTAL RESULTS AND DISCUSSIONExtrapolated predictions of creep results

The extrapolated creep displacements at 600 days are 0.26 and 3.2 mm (0.010 and 0.124 in.) for Systems C and F, respectively. ASTM E1512 recommends comparing creep displacement to ultimate displacement during a tensile test; however, ICC-ES Acceptance Criteria 58 (AC58)4 is more explicit in its creep displacement limits. AC584 specifies displacement at 600 days to be below the lesser of displace-ment at ultimate load or 3.05 mm (0.12 in.). Taking 3 mm (0.12 in.) as the limit, it can be seen that System C meets the requirements of AC584 but that System F slightly exceeds them. The disparity between the extrapolated creep results of the two systems is significant and these results provide another example of the large disparities in the long-term performance of different ASTM C881/C881M systems. The difference in the performance of the two epoxy systems can be seen in Fig. 5, which compares the specimens after the conclusion of creep testing. An arrow indicates the location of visible, permanent creep displacement on the System F specimen.

Creep comparison to other types of reinforcing bar couplers

Epoxy-bonded reinforcing bar couplers are similar to cementitiously grouted reinforcing bar couplers; the only difference is the type of material used to grout the reinforcing bars into a steel sleeve. In one published creep test, cementi-tiously grouted reinforcing bar couplers were found to have creep displacements extrapolated to 600 days of approxi-mately 0.75 mm (0.030 in.).3 This value is less than the maximum displacement recommended by AC58 of 3 mm (0.12 in.) and also matches the performance of System C evaluated in this research.

FURTHER RESEARCHBecause epoxy degrades more quickly in moister and/or

higher-temperature environments, additional research to inves-tigate the creep performance of epoxy-bonded reinforcing bar couplers at a range of temperatures and moist environments should be undertaken. To understand the potential impact of heat and moisture on the creep performance of epoxy-bonded reinforcing bar couplers, the environmental condi-tions that can be expected during service inside a reinforced concrete structure should be used.

CONCLUSIONSBased on the results of this experimental investigation of

creep loading, the following conclusions are drawn:1. Epoxy-bonded reinforcing bar couplers do creep

measurably due to the creep of the epoxy grout under sustained load. This phenomenon is similar to the creep behavior of epoxy when used for anchorage to concrete.

2. The extrapolated creep performance of epoxy-bonded reinforcing bar couplers can vary widely, depending on which epoxy system is used to assemble them.

3. Epoxy-bonded reinforcing bar couplers, if assembled with a well-suited epoxy system, have adequate creep performance in dry, room-temperature conditions.

4. Epoxy-bonded reinforcing bar couplers, if assembled with a well-suited epoxy system, have similar creep perfor-mance to other types of nonadhesive, grouted reinforcing bar couplers.

ACKNOWLEDGMENTSThe author wishes to express his gratitude and sincere appreciation to

the California Department of Transportation for financing this research, V. Karbhari for overseeing the research, and A. Pridmore and the staff of the UCSD Powell and SRMD Laboratories for their assistance.

REFERENCES1. ASTM C881/C881M-02, “Standard Specification for Epoxy-

Resin-Base Bonding Systems for Concrete,” ASTM International, West Conshohocken, PA, 2002, 6 pp.

2. ASTM E1512-01(2007), “Standard Test Methods for Testing Bond Performance of Bonded Anchors,” ASTM International, West Conshohocken, PA, 2001, 5 pp.

3. Jansson, P. O., “Evaluation of Grout-Filled Mechanical Splices for Precast Concrete Construction,” Report No. TI-2094, Michigan Depart-ment of Transportation, Construction and Technology Division, Lansing, MI, 2008, 68 pp.

4. ICC-ES Acceptance Criteria 58 (AC58), “Acceptance Criteria for Adhesive Anchors in Concrete and Masonry Elements,” International Congress of Building Officials Evaluation Services (ICBO ES), 2001, 17 pp.

Fig. 5—Photograph of reinforcing bar coupler specimens after conclusion of testing.


Title no. 109-M48


ACI Materials Journal, V. 109, No. 5, September-October 2012.MS No. M-2011-033.R3 received December 6, 2011, and reviewed under Institute


Effect of Na2SiO3/NaOH Ratios and NaOH Molarities on Compressive Strength of Fly-Ash-Based Geopolymerby A. M. Mustafa Al Bakri, H. Kamarudin, M. Bnhussain, A. R. Rafiza, and Y. Zarina

and moldable paste) and stored at mild temperatures (T < 100°C [212°F]) for a short period of time to produce a mate-rial with good binding properties. At the end of this process, an amorphous alkaline alumino-silicate gel is formed as the main reaction product. In addition, Na-herschelite-type zeolites and hydroxysodalite are formed as secondary reac-tion products.4-6

The most-used alkaline activators are a mixture of sodium or potassium hydroxide (NaOH or KOH) with sodium water glass (nSiO2Na2O) or potassium water glass (nSiO2K2O).1,7-9 One of the factors that influences the compressive strength of geopolymer is the Na2SiO3/NaOH ratio.10,11 Rattanasak and Chindaprasirt12 concluded that the use of an Na2SiO3/NaOH ratio of 1.0 produced a product with a compressive strength as high as 70 MPa (10.15 ksi). A study conducted by Hardjito et al.10 showed that the use of an Na2SiO3/NaOH ratio of 2.5 gave the highest compres-sive strength of 56.8 MPa (8.24 ksi), whereas a ratio of 0.4 resulted in a lower compressive strength of 17.3 MPa (2.51 ksi).

The concentrations of NaOH solution that can be used are in the range of 8 to 16 M.8 Some researchers7,12,13 have studied the effects of different molarities of NaOH on the geopolymer. Puertas et al.13 stated that, at 28 days of reac-tion, a mixture of equal parts FA and slag activated with 10 M NaOH and cured at 25°C (77°F) develops a compressive strength of approximately 50 MPa (7.25 ksi). Rattanasak and Chindaprasirt12 concluded that a geopolymer mortar strength of up to 70 MPa (10.15 ksi) is obtained when the mixture is formulated with 10 M NaOH and an Na2SiO3/NaOH ratio of 1.0. Palomo et al.7 reported that a 12 M activator concentra-tion leads to better results than an 18 M concentration.

In this study, the effects of various Na2SiO3/NaOH ratios and NaOH molarities on FA geopolymer paste were studied. Six different Na2SiO3/NaOH ratios (0.5, 1.0, 1.5, 2.0, 2.5, and 3.0) and six different NaOH molarities (6, 8, 10, 12, 14, and 16 M) were used in this study. The geopolymer properties, such as compressive strength, water absorption, porosity, and density, were used as indicators to prove that the geopolymer has similar properties to PC.

RESEARCH SIGNIFICANCEThe Na2SiO3 and NaOH solution requires different mass

proportions with different FAs to obtain high compressive strength. Most of the studies on geopolymer concentrated

Carbon dioxide (CO2) emissions from the production of 1 ton (2204.62 lb) of cement vary between 0.05 and 0.13 tons (110.23 and 286.60 lb). It is important to reduce CO2 emissions by the greater use of substitutes for portland cement (PC), such as fly ash (FA), clay, and other geo-based materials. This paper studies the processing of geopolymers using FA and alkaline activators. The factors that influence the early-age compressive strength, such as the sodium hydroxide (NaOH) molarity and Na2SiO3/NaOH ratios, were studied. Sodium hydroxide and sodium silicate solutions were used as alkaline activators. The geopolymer paste samples were cured at 70°C (158°F) for 1 day and kept at room temperature until testing (the seventh day). The compressive strength was measured after 7 days. The results show that the geopolymer paste with a combination of an Na2SiO3/NaOH ratio of 2.5 and a 12 M NaOH concentration produces the highest compressive strength. The density obtained for geopolymer for PC is in the range of 1760 to 1855 kg/m3 (0.064 to 0.067 lb/in.3). The porosity of the geopolymer was in the range of 12.16 to 26.19%, and the water absorption was in the range of 5.03 to 8.13%. The results of scanning electron microscopy (SEM) indicated that the samples with a denser matrix and less unreacted FA contributed to the maximum compressive strength. In the X-ray diffraction (XRD) patterns, the intensity of quartz content at 12 M was highly detected compared to the 6 and 10 M solutions.

Keywords: alkaline activation; compressive strength; geopolymer; Na2SiO3/NaOH ratio; NaOH molarity; scanning electron microscopy; X-ray diffraction.

INTRODUCTIONThe term “geopolymer” was first applied by Davido-

vits1 to alkali alumino-silicate binders formed by the alkali-silicate activation of alumino-silicate materials. Geopolymers (green polymeric concrete) are amorphous to the semi-crystalline equivalent of certain zeolitic mate-rials with excellent properties, such as high fire and erosion resistances, as well as high strength. Recent works2 have shown that the addition of moderate amounts of minerals to a geopolymer can yield significant improvements in the geopolymer’s structure and properties.

The alkaline liquid could be used to react with the silicon (Si) and aluminum (Al) in a source material of natural minerals or in by-product materials, such as fly ash (FA) and rice husk ash, to produce binders.1 The alkaline activation of materials can be defined as a chemical process that provides a rapid change of specific structures—partial or completely amorphous—into compact cemented frameworks.3 The alkali activation of FA is a process that differs widely from portland cement (PC) hydration and is very similar to the chemistry involved in the synthesis of large groups of zeolites.4 Some researchers5,6 have described the alkali activation of FA (AAFA) as a physicochemical process in which this powdery solid is mixed with a concentrated alkali solution (in a suitable proportion to produce a workable


A. M. Mustafa Al Bakri is a Senior Lecturer and PhD Candidate at Universiti Malaysia Perlis (UniMAP), Perlis, Malaysia. He received his BS in civil engineering and his MS in material engineering from Universiti Sains Malaysia (USM), Penang, Malaysia. His research interests include green and construction materials.

H. Kamarudin is a Vice Chancellor at UniMAP. He received his BS, MS, and PhD in chemistry from USM. His research interests include chemistry reaction and sustain-able material.

M. Bnhussain is a Director of the Program of Advanced Material at the King Abdu-laziz City for Science and Technology, Riyadh, Saudi Arabia. He received his BS in civil engineering from King Abdulaziz University, Jeddah, Saudi Arabia, and his PhD in civil engineering materials from the University of Leeds, Leeds, UK. His research interests include construction materials, including green polymeric concrete.

A. R. Rafiza is a Researcher at UniMAP. She received her BS and MS in civil engi-neering (structural engineering) from USM. Her research interests include seismic modeling, structural analysis, and green polymeric concrete.

Y. Zarina is a Researcher at UniMAP. She received her BS and MS in civil engi-neering (structural engineering) from USM. Her research interests include seismic modeling, structural analysis, and green polymeric concrete.

Table 1—Chemical composition of FA

Chemical composition Percentage, %

SiO2 52.11

Al2O3 23.59

Fe2O3 7.39

TiO2 0.88

CaO 2.61

MgO 0.78

Na2O 0.42

K2O 0.80

P2O5 1.31

SO3 0.49

MnO 0.03

Table 2—Mixture design details for various ratio of Na2SiO3/NaOH

FA/alkaline activator ratio

Na2SiO3/NaOH ratio

FA,lb (g)

Na2SiO3,lb (g)

NaOH,lb (g)

2.5

0.5

1.33 (605)

0.18 (80) 0.35 (160)

1.0 0.26 (120) 0.26 (120)

1.5 0.32 (145) 0.21 (95)

2.0 0.35 (160) 0.18 (80)

2.5 0.37 (170) 0.15 (70)

3.0 0.40 (180) 0.13 (60)

on only two different molarities of NaOH. This study deals with more details on different NaOH molarities (6, 8, 10, 12, 14, and 16 M) of geopolymer pastes. The research data presented in this paper are useful to understand the effect of various Na2SiO3/NaOH ratios and different molarities on the geopolymer, which influence the compressive strength results. The compressive strength of specimens decreases with increasing porosity and water absorption. The scanning electron microscopy (SEM) and X-ray diffraction (XRD) tests are also important to understand the microstructural characteristics and phases involved in geopolymer.

EXPERIMENTAL PROCEDUREMaterials

In this research, low-calcium Class F dry FA14 obtained from the Sultan Abdul Aziz Power Station in Kapar, Selangor, Malaysia, was used as a base material to make the geopolymers. The chemical composition of FA is shown in Table 1. The table shows that this FA consists of a high composition of silicon and aluminum oxide of 75.7%.

The mixture of sodium silicate (Na2SiO3) and sodium hydroxide (NaOH) was used as an alkaline activator in this study. NaOH in pellet form with 97% purity8,15,16 and Na2SiO3 consisting of Na2O = 9.4%, SiO2 = 30.1%, and H2O = 60.5% (with a weight ratio of SiO2/Na2O of 3.20 to 3.30 and a specific gravity of 20°C [68°F] = 1.4 g/cm3 [0.05 lb/in.3]) were used in this study.

Mixing methodThe ratio of FA to alkaline activator was 2.5 and was

kept fixed for all mixtures. The use of this ratio is due to the work of Hardjito et al.,10,17 which states that a ratio of FA to alkaline activator of 2.5 produces the highest compres-sive strength on the 28th day of testing. In this study, various Na2SiO3/NaOH ratios (0.5, 1.0, 1.5, 2.0, 2.5, and 3.0) were used to determine the highest compressive strength. It should be noted that Na2SiO3 is a quick-setting chemical and binding material that requires a different combination of proportions with NaOH molarities. The NaOH molarity was kept constant at 10 M. The total mass of each dry and solutions material used is shown in Table 2. The best ratio of the Na2SiO3/NaOH result (highest compressive strength) was further used in this study.

To prepare the NaOH solution, NaOH pellets were dissolved in 1 L (0.26 gal.) of distilled water in a volumetric flask for six different NaOH concentrations (6, 8, 10, 12, 14, and 16 M) with different masses of NaOH, as shown in Table 3. The best mixture design (the FA/alkaline activator and Na2SiO3/NaOH ratios were fixed as 2.5 as their highest compressive strength result obtained previously) was used to determine the best NaOH molarity solution. The total mass of each of the dry and solutions materials used was kept constant for the geopolymer paste for all samples (6, 8, 10, 12, 14, and 16 M) with different molarities.

An alkaline activator consisting of a combination of NaOH and Na2SiO3 was prepared just before mixing with FA to ensure the reactivity of the solution. The addition of sodium silicate is to enhance the geopolymerization process.18 The FA and alkaline activator were mixed together in the mixer until a homogeneous paste was achieved. This mixing process could be handled for up to 10 minutes for each mixture with different NaOH molarities, as shown in Fig. 1.

Table 3—Details of preparing NaOH solutions

NaOH molarity, MMasses of NaOH pellets dissolved in

1 L (0.26 gal.) of distilled water, lb (g)

6 0.53 (240)

8 0.71 (320)

10 0.88 (400)

12 1.06 (480)

14 1.23 (560)

16 1.41 (640)


Casting and curingThe geopolymer paste was placed in a 50 x 50 x 50 mm

(1.97 x 1.97 x 1.97 in.) cube mold and cured in an oven for 1 day1 at 70°C (158°F).1,19 After the samples were cured in an oven, the molds were removed from the furnace and left to cool to room temperature before demolding.1 The samples were then left to room temperature until they were loaded in compression at the seventh day.1

TestingThe compressive strength test was performed on

geopolymer paste samples in accordance with BS 1881-116:198320 using a mechanical testing machine to obtain the ultimate strength of the geopolymer. The samples were loaded with 50.00 kN (11.24 kips) and the speed rate of loading was 5.00 mm/min (0.20 in./min). The loading pace rate was 0.1 kN/s (22.48 lb/s. The reported compressive strength values are an average of three samples for each ratio.

The sample densities were determined by the mass and volumes of the cubes in accordance with BS 1881-114:1983.21 The results of densities are taken as an average of three samples for each ratio.

The water absorption test was performed in accordance with ASTM C140 to determine the porosity of the samples. The sample masses were measured before and after immer-sion in water. The difference in weight was calculated to determine the water absorption of the samples, as shown in Eq. (1).

Water Absorption 100S D

D

M MM−

= • (1)

where MS is saturated mass, units; and MD is dry mass, units.A scanning electron microscope was used to reveal the

microstructure of the geopolymer paste. The test was carried out using secondary and backscattered electron detectors.

XRD patterns were performed using an X-ray diffractom-eter. The XRD test was held for phase analysis of the orig-inal FA and to investigate the crystallinity of the geopolymer samples that gave high compressive strength. The samples were prepared in powder form. For the prepared geopolymer samples, the samples were first cut into 0.5 mm (1.97 in.) thick slices and then ground into powder form as required.

EXPERIMENTAL RESULTS AND DISCUSSIONCompressive strength

The Na2SiO3/NaOH ratio and NaOH molarity affects the compressive strength of the geopolymer. The compressive strengths for different Na2SiO3/NaOH ratios are shown in Fig. 2. The highest compressive strengths of 57.00 MPa (8.27 ksi) were observed at an Na2SiO3/NaOH ratio of 2.5, which is 18% higher than an Na2SiO3/NaOH ratio of 3.0 on the seventh day of testing. Hardjito et al.10 and Sathia et al.22 stated that compressive strength increases as FA content and concentration of the activator solution increase. This is due to the increase in the sodium oxide content, which is mainly required for the geopolymerization reaction. The compressive strength of the product for an Na2SiO3/NaOH ratio of 3.0 was low, however, which could be due to the excess OH– concentration in the mixtures.18 Furthermore, the excess sodium content can form sodium carbonate by atmo-spheric carbonation and this may disrupt the polymerization

process.23 The lowest compressive strength was found at an Na2SiO3/NaOH ratio of 0.5 with 40.00 MPa (5.80 ksi).

The compressive strength results for various NaOH molar-ities on the seventh day of testing are shown in Fig. 3. For the seventh day of testing, the 12 M NaOH samples produced the highest compressive strengths of 94.59 MPa (13.72 ksi). This is due to the increase of Na ions in the system, which was important for the geopolymerization because Na ions were used to balance the charges and formed the alumino-silicate networks as the binder in the mixture.24 At a low

Fig. 1—Mixture of FA with alkaline activator.

Fig. 2—Compressive strength of various Na2SiO3/NaOH ratios.

Fig. 3—Compressive strength of various NaOH molarities.


which reduced the compressive strength of the sample. The amount of liquid in the systems affects the saturation rate of the ionic species and the strength of the geopolymer. The geopolymer microstructures with different NaOH molarities are shown in Fig. 5(a) through (f).

As the NaOH molarity increases from 6 to 16 M, the microstructure of the resultant geopolymer contains a smaller proportion of unreacted FA microspheres. As can be seen from Fig. 5(a) and (b), a large proportion of FA still did not completely dissolve. Figure 5(d) shows the least unreacted FA for the alkaline activator that gave the highest compressive strength of 94.59 MPa (13.72 ksi) on the seventh testing day. This suggests that the dissolution of silica and alumina in the geopolymerization process that formed the alumino-silicate gel in the 12 M NaOH sample was higher, contributing to the stability of the geopolymer during the hardening process and giving a high compressive strength of geopolymer.29 The pores are indicated on the figures by arrows, and cracks are also found in the matrix (Fig. 5(a), (b), (c), (e), and (f)), which would limit the binding capacity and lead to a lower compressive strength.

XRD patternThe results of the XRD of FA and geopolymer pastes with

6, 10, and 12 M NaOH concentrations are shown in Fig. 6. The original FA and the 6, 10, and 12 M NaOH molarity pastes had a similar diffraction pattern and did not signifi-cantly alter the degree of amorphous and crystallization of FA. For 12 M NaOH molarity, the XRD pattern showed that the intensity of quartz content was highly detected at 2q = 26.5 degrees compared to the 6 and 10 M solutions. This also indicated that the new crystalline phases were detected in the geopolymer paste and that the 12 M NaOH solution contains the highest amount of crystalline and had a higher compres-sive strength compared to FA. Alvarez-Ayuso et al.30 stated that the increase in the crystalline product increased the compressive strength of the geopolymer. The formation of crystallines in the samples—studied by quantitative XRD—depended strongly on the NaOH concentration. The crystal-lization rate increased with increasing NaOH content and the proportion of the crystalline phase gradually increased with a longer curing time.31

The obtained results suggest that the composition of the alumino-silicate gel formed by the reaction between FA and

NaOH molarity, the geopolymerization is low due to the low concentration of base and, hence, less leaching of silica and alumina from the source material.25 The lowest compressive strength was found for the 6 M NaOH solu-tion with 40.00 MPa (5.8 ksi). After 12 M of NaOH solu-tion, the compressive strength decreased. The high viscosity hinders the leaching of the silica and alumina, resulting in a lesser degree of geopolymerization26 as compared to that of the 12 M NaOH paste. Palomo et al.27 also found that a 12 M NaOH solution produced better results than the corresponding 18 M activator.

The compressive strength of PC paste was in the range of 17 to 20 MPa (2.47 to 2.90 ksi) at 28 days of testing; however, geopolymer paste can achieve a better performance of compressive strength of 94.59 MPa (13.72 ksi) at 7 days of testing. This clearly shows that geopolymer paste can achieve a higher compressive strength than PC paste.

Density, porosity, and water absorptionThe densities of the geopolymer samples are in the range

of 1760 to 1855 kg/m3 (0.064 to 0.067 lb/in.3) for 7 days of testing. Higher alkali contents in the mixture yield better reactivity with the FA, resulting in a denser microstruc-ture.28 The density of normal PC paste is 1750 to 2400 kg/m3 (0.063 to 0.087 lb/in.3). Because the density obtained from the geopolymer samples is in this range, the samples possess the same properties as PC paste.

The porosity of the geopolymer was in the range of 12.16 to 26.19%, and the paste specimen produced water absorption in the range of 5.03 to 8.13% for the seventh day of testing. According to Thokchom et al.,28 the compressive strength of specimens decreases with increasing porosity and water absorption.

SEM analysis for geopolymer pasteThe microstructure of FA-based geopolymer for different

mixture designs was observed with SEM, as shown in Fig. 4(a) and (b). It showed that the materials are hetero-geneous, with partially reacted and unreacted FAs existing on the dense, gel-like matrix geopolymer. The sample with the FA/alkaline activator and an Na2SiO3/NaOH ratio of 2.5 (Fig. 4(a)) showed a more dense matrix and less unreacted FA, which contributed to a maximum compressive strength of 8.27 ksi (57 MPa). Figure 4(b) shows the microcracks that exist on the sample with an FA/alkaline activator ratio of 2.5,

Fig. 4—SEM pictures of geopolymer paste with various mixture designs: (a) FA/alka-line activator of 2.5 and Na2SiO3/NaOH of 2.5; and (b) FA/alkaline activator of 2.5 and Na2SiO3/NaOH of 3.0.


the alkaline activator is variable and depends on the reac-tivity and the type and concentration of the activators.32

CONCLUSIONSBased on the experimental work reported in this paper, it

can be concluded that the Na2SiO3/NaOH ratios and NaOH molarities affect the compressive strength of FA-based geopolymer. The Na2SiO3/NaOH ratio of 2.5 contributed to the high compressive strength of 57.00 MPa (8.27 ksi). The highest NaOH molarity does not necessarily give the highest compressive strength. The FA-based geopolymer with 12 M NaOH showed excellent results, including a high compres-sive strength of up to 94.59 MPa (13.72 ksi) on the seventh testing day. This was proven by the XRD results, which show that the intensity of the quartz content at 12 M was highly detected and contributed to the highest compressive strength compared to the 6 and 10 M solutions. The density obtained for the geopolymer (1760 to 1855 kg/m3 [0.064 to 0.067 lb/in.3]) was in the range for PC of 1750 to 2400 kg/m3 (0.063 to 0.087 lb/in.3). The samples with a denser matrix and less unreacted FA contributed to the maximum compres-

Fig. 5—SEM image of geopolymer with: (a) 6 M; (b) 8 M; (c) 10 M; (d) 12 M; (e) 14 M; and (f) 16 M of NaOH solution.

Fig. 6—XRD patterns of FA and geopolymer paste samples with 6, 10, and 12 M NaOH concentrations.


15. Buchwald, A., and Schulz, M., “Alkali-Activated Binders by Use of Industrial By-Products,” Cement and Concrete Research, V. 35, No. 5, May 2005, pp. 968-973.

16. Wallah, S. E., “Drying Shrinkage of Heat-Cured FA-Based Geopolymer Concrete,” Modern Applied Science, V. 3, No. 12, Dec. 2009, pp. 12-21.

17. Hardjito, D.; Cheak, C. C.; and Lee Ing, C. H., “Strength and Setting Time of Low Calcium FA-Based Geopolymer Mortar,” Modern Applied Science, V. 2, No. 4, July 2008, pp. 3-11.

18. Xu, H., and Deventer, J., “The Geopolymerisation of Alumino-Sili-cate Minerals,” International Journal of Mineral Processing, V. 59, No. 3, 2000, pp. 247-266.

19. Van Jaarsveld, J. G. S.; Van Deventer, J. S. J.; and Lukey, G. C. C., “The Effect of Composition and Temperature on the Properties of Fly Ash and Kaolinite-Based Geopolymers,” Chemical Engineering Journal, V. 4001, 2002, pp. 1-11.

20. BS 1881-116:1983, “Testing Concrete. Method for Determination of Compressive Strength of Concrete Cubes,” British Standards Institution, London, UK, 1983, 8 pp.

21. BS 1881-114:1983, “Testing Concrete. Methods for Determination of Density of Hardened Concrete,” British Standards Institution, London, UK, 1983, 8 pp.

22. Sathia, R.; Ganesh Babu, K.; and Santhanam, M., “Durability Study of Low Calcium FA Geopolymer Concrete,” The 3rd ACF International Conference ACF/VCA, 2008, pp. 1153-1159.

23. Barbosa, V. F. F.; Mackenzie, K. J. D.; and Thaumaturgo, C., “Synthesis and Characterisation of Materials Based on Inorganic Poly-mers of Alumina and Silica: Sodium Polysialate Polymers,” International Journal of Inorganic Materials, V. 2, 2000, pp. 309-317.

24. Sathonsaowaphak, A.; Chindaprasirt, P.; and Pimraksa, K., “Work-ability and Strength of Lignite Bottom Ash Geopolymer Mortar,” Journal of Hazardous Materials, V. 168, No. 1, Aug. 2009, pp. 44-50.

25. Alonso, S., and Palomo, A., “Alkaline Activation of Metakaolin and Calcium Hydroxide Mixtures: Influence of Temperature, Activator Concen-tration and Solid Ratio,” Materials Letters, V. 47, No. 1-2, 2001, pp. 55-62.

26. Chindaprasirt, P.; Jaturapitakkul, C.; Chalee, W.; and Rattanasak, U., “Comparative Study on the Characteristic of FA and Bottom Ash Geopolymer,” Waste Management, V. 29, No. 2, Feb. 2009, pp. 539-543.

27. Palomo, A.; Blanco, M.; Granizo, M.; Puertas, F.; Vazquez, T.; and Grutzeck, M., “Chemical Stability of Cementitious Materials Based on Metakaolin,” Cement and Concrete Research, V. 29, No. 7, July 1999, pp. 997-1004.

28. Thokchom, S.; Ghosh, P.; and Ghosh, S., “Effect of Water Absorp-tion, Porosity, and Sorptivity on Durability of Geopolymer Mortars,” ARPN Journal of Engineering and Applied Sciences, V. 4, No. 7, Sept. 2009, pp. 28-32.

29. Bakharev, T., “Resistance of Geopolymer Materials to Acid Attack,” Cement and Concrete Research, V. 35, No. 4, 2005, pp. 658-670.

30. Alvarez-Ayuso, E.; Querol, X.; Alastuey, A.; Moreno, N.; Izqui-erdo, M.; Font, O.; Moreno, T.; Ramonich, E. V.; Diez, S.; and Barra, M., “Environmental, Physical and Structural Characterisation of Geopolymer Matrixes Synthesised from Coal (Co-)Combustion FAes,” Journal of Hazardous Materials, V. 154, 2008, pp. 175-183.

31. Zhang, B.; MacKenzie, K. J. D.; and Brown, I. W. M., “Crystalline Phase Formation in Metakaolinite Geopolymers Activated with NaOH and Sodium Silicate,” Journal of Materials Science, V. 44, 2009, pp. 4668-4676.

32. Duchesne, J.; Duong, L.; Bostrom, T.; and Frost, R., “Microstructure Study of Early In Situ Reaction of FA Geopolymer Observed by Environ-mental Scanning Electron Microscopy (ESEM),” Waste and Biomass Valo-rization, V. 1, No. 3, 2010, pp. 367-377.

sive strength. A higher dissolution of silica and alumina in the geopolymerization process that forms alumino-silicate gel contributes to the higher compressive strength of the geopolymer; however, different FAs from other countries may need different ratios to achieve high compressive strength. FA-based geopolymer has excellent properties due to the very high compressive strength obtained in this study. Further studies need to be conducted to find the best mixture design to achieve the highest compressive strength of FA-based geopolymer concrete.

ACKNOWLEDGMENTSA grant from the King Abdulaziz City for Science and Technology

(KACST) for this research project is gratefully acknowledged.

REFERENCES1. Davidovits, J., “High-Alkali Cements for 21st Century Concretes,”

Concrete Technology: Past, Present, and Future, SP-144, P. K. Mehta, ed., American Concrete Institute, Farmington Hills, MI, 1994, pp. 383-397.

2. Hu, M.; Zhu, X.; and Long, F., “Alkali-Activated FA-Based Geopoly-mers with Zeolite or Bentonite as Additives,” Journal of Cement and Concrete Composites, V. 31, No. 10, Nov. 2009, pp. 762-768.

3. Fernández-Jiménez, A., and Palomo, A., “Characterisation of FAes. Potential Reactivity as Alkaline Cements,” Fuel, V. 82, No. 18, Dec. 2003, pp. 2259-2265.

4. Criado, M.; Palomo, A.; and Fernández-Jiménez, A., “Alkali Activa-tion of FAes. Part 1: Effect of Curing Conditions on the Carbonation of the Reaction Products,” Fuel, V. 84, No. 16, Nov. 2005, pp. 2048-2054.

5. Fernández-Jiménez, A.; de la Torre, A. G.; Palomo, A.; Lopez-Olmo, G.; Alonso, M. M.; and Aranda, M. A. G., “Quantitative Determina-tion of Phases in the Alkaline Activation of FA. Part II: Degree of Reaction,” Fuel, V. 85, No. 14-15, Oct. 2006, pp. 1960-1969.

6. Fernández-Jiménez, A., and Palomo, A., “Composition and Micro-structure of Alkali Activated FA Binder: Effect of the Activator,” Cement and Concrete Research, V. 35, No. 10, Oct. 2005, pp. 1984-1992.

7. Palomo, A.; Grutzek, M. W.; and Blanco, M. T., “Alkali-Activated FAes: A Cement for the Future,” Cement and Concrete Research, V. 29, No. 8, Aug. 1999, pp. 1323-1329.

8. Rangan, B. V., Concrete Construction Engineering Handbook, Taylor and Francis Group, LLC, London, UK, 2008, pp. 1-19.

9. Pacheco-Torgal, F.; Castro-Gomes, J.; and Jalali, S., “Alkali-Activated Binders: A Review. Part 2. About Materials and Binders Manufacture,” Journal of Construction and Building Material, V. 22, No. 7, July 2008, pp. 1315-1322.

10. Hardjito, D.; Wallah, S. E.; Sumajouw, D. M. J.; and Rangan, B. V., “On the Development of FA-Based Geopolymer Concrete,” ACI Materials Journal, V. 101, No. 6, Nov.-Dec. 2004, pp. 467-472.

11. Mustafa Al Bakri, A. M.; Kamarudin, H.; BnHussain, M.; Khairul Nizar, I.; Rafiza, A. R.; and Zarina, Y., “Microstructure of Different NaOH Molarity of Fly Ash-Based Green Polymeric Cement,” Journal of Engi-neering and Technology Research, V. 3, No. 2, 2011, pp. 44-49.

12. Rattanasak, U., and Chindaprasirt, P., “Influence of NaOH Solution on the Synthesis of FA Geopolymer,” Minerals Engineering, V. 22, No. 12, Oct. 2009, pp. 1073-1078.

13. Puertas, F.; Martinez-Ramirez, S.; Alonso, S.; and Vazquez, T., “Alkali-Activated FA/Slag Cement: Strength Behaviour and Hydration Products,” Cement and Concrete Research, V. 30, No. 10, Oct. 2000, pp. 1625-1632.

14. Chindaprasirt, P.; Chareerat, T.; and Sirivivatnanon, V., “Work-ability and Strength of Coarse High Calcium FA Geopolymer,” Cement and Concrete Composites, V. 29, No. 3, Mar. 2007, pp. 224-229.


Title no. 109-M49


ACI Materials Journal, V. 109, No. 5, September-October 2012.MS No. M-2011-042.R3 received January 3, 2012, and reviewed under Institute


Effect of Using Mortar Interface and Overlays on Masonry Behavior by Using Taguchi Methodby Mariam Farouk Ghazy

to out-of-plane and in-plane vertical and lateral stresses was investigated.4 A significant strength increase was observed for all strengthened specimens.

Experimental investigations that studied the rehabilita-tion of masonry walls with reinforced mortar overlays were carried out.5,6 The general conclusion was that the applica-tion of mortar overlays is a powerful rehabilitation technique for masonry constructions.

The results of a series of axial compression tests on concrete block wallets coated with cement mortar overlays were presented. Different types of mortars and combina-tions with steel welded meshes and fibers were tested.6,7 The main conclusion was that the application of mortar overlays increases the wall strength, but not in a uniform manner; the strengthening efficiency of wallets loaded in axial compres-sion is not proportional to the overlay mortar strength because it can be affected by the failure mechanisms of the wall. Steel mesh-reinforced overlays, in combination with high-strength mortar, show better efficiency because the steel mesh mitigates the damage effects in the block wall and in the overlays themselves.7

A relationship between the masonry prism compressive strength and bond strength was obtained.8 The results clearly indicate that an increase in bond strength, while keeping the mortar strength constant, leads to an increase in the compres-sive strength of masonry.

The various influences on masonry behavior caused by brick-mortar interface properties were investigated.9 The brick-mortar interface is characterized by a central bonded area of variable size and shape surrounded by fissures near the masonry surface. The interface is usually the weakest spot during bending and shear. At the fissure tip (10 to 15 mm [0.39 to 0.59 in.] deep from the surface), the bricks split, showing that high strain levels develop around the central brick-mortar contact area. In stack-bonded masonry prisms, the fissures close before the bricks notice-ably deform. Only after closing of the horizontal fissures, which occurs at considerable loads, do the outer sides of the bricks become stressed.

The behavior of lime-based renders on the masonry walls made from solid clay bricks and parallel tests of the characteristics of fresh and hardened lime-based mortars was studied.10 The tests were carried out on five different lime-based mortar mixtures. The results show that in cases when higher ductility of the hardened lime-based renders

The effect and optimization of using different types of mortar for both interface and overlays on masonry behavior was investi-gated by using the Taguchi method. The experimental studies were conducted under varying types of mortar. The mortar was rein-forced with a polypropylene (PP) fiber with a volume fraction (0, 1, and 2%) by volume of mortar. An orthogonal array (OA), the signal-to-noise ratio (S/N), and the analysis of variance (ANOVA) were employed to study the performance characteristics of the masonry prisms and walls.

The conclusion revealed that the number of overlays and the type of interface mortar were the most influential factors on the masonry prism’s compressive strength and flexural bond strength, respectively. Moreover, the application of mortar over-lays increases the wall compressive strength.

Keywords: compressive strength; flexural bond strength; masonry walls; overlay; polypropylene fibers; signal-to-noise ratio; Taguchi method.

INTRODUCTIONMasonry—specifically brick and mortar—requires much

more than knowledge of the brick or mortar individually to fully understand its properties. Why would two materials of very different strengths combine to form a composite system that displays a yield strength intermediate to both and not that of the lower-strength material? This materials interaction is being reviewed at the University of Colorado, Boulder.1 Recent advances in masonry technology brought new materials, building techniques, and rational methods of structural analysis; however, the structural behavior of masonry walls is still a complex matter. A concrete masonry wall is made of at least two different materials that are assembled under diverse conditions of execution and quality control. If the masonry wall is coated on both sides with cement mortar overlays, these overlays become part of the composite element. In this case, the wall can be seen as a sandwich panel where the overlays are the covering sheets and the concrete masonry wall is the core.2

The use of fiber-reinforced polymers (FRPs) as external wraps for the structural rehabilitation of buildings and bridges has taken tremendous strides forward over the past decade. In particular, extensive use of this structural rehabili-tation system has been made in the area of masonry struc-tures.2,3 The glass fiber-reinforced polymer (GFRP) wrap system provided adequate tension reinforcement to develop the compressive strength of the masonry, thus substantiating its utility in this application.

Epoxy-bonding a thin layer of composite materials to the exterior surfaces of unreinforced masonry (URM) walls forces the individual brick or block elements to act as an integrated system. The high tensile strength of composite materials can be used to significantly increase the shear and flexural capacity of URM walls. The application of GFRP laminates in strengthening the concrete block walls subjected


has been made to investigate the effect of using mortar interface and overlays on masonry behavior by using the Taguchi11 method. Furthermore, the analysis of variance (ANOVA) was used to discuss the relative importance of all control factors.

RESEARCH SIGNIFICANCEThe effect and optimization of using different types of

mortar for both interface and overlays on masonry behavior was investigated in this study by using the Taguchi11 method as a new technique in experimental design. Taguchi’s11 method of experimental design was used in this study to provide a simple, efficient, and systematic approach for the optimi-zation of experimental designs. The experimental studies were conducted under varying types of mortars to study the masonry behavior. Three basic control factors were taken into consideration: the type of mortar interface (A), the type of mortar overlays (B), and the number of overlays (C).

EXPERIMENTAL WORKMaterials and means

The constituent materials used in this study were locally available materials specified by the following:

1. Brick units: Perforated shale brick units (10 vertical holes) were obtained from various manufacturers to select the brick with the best performance (with a higher compres-sive strength and low absorption). The absorption proper-ties of the brick may affect the mortar structure and, conse-quently, the mechanical behavior.9 The mechanical prop-erties of the brick units are given in Table 1, according to ES 4763/200513 and ECP 204-2005.14 A hydraulic testing machine with a total capacity of 348.2 kips (1550 kN) was used to test brick units. No. 3 brick was used to complete the experimental program, and the area of the holes was less than 25% of the total surface area of the brick units. Thus, the loading area was taken as equal to the gross area.

2. Cement: Grade 4700 psi (32.5 N) ordinary portland cement was used in this investigation. Cements conform to ES 4756-1/2007.15

3. Fine aggregates: Medium well-graded sand with a fineness modulus of 2.2 and 2.5 was used for mortar and concrete, respectively.

4. Coarse aggregates: Natural well-graded gravel with a maximum nominal size of 0.787 in. (20 mm) was used for casting the concrete beam. It included a combination of round and angular particles. The surface of the particles was more or less smooth. The fine and coarse aggregates conformed to ES 1109/2002.16

5. Chemical admixtures: A high-range water-reducing admixture (HRWRA) was used in fiber-reinforced mortar mixtures to keep a plastic consistency of mortar that satisfies the requirements of ASTM C494/C494M-99a17 Type F and BS 5075-3:198518 for HRWRA. Its dosage ranged between 0.6 and 2.5% of cement weight, as given by the manufacturer.

6. Fibers: The polypropylene (PP) fibers used in this inves-tigation are commercially available. The length of the fibers was approximately 0.75 in. (19 mm) and the equivalent diameter was 0.0016 in. (0.04 mm).

Specimen preparation and testingThe mortar mixtures were weighed and mixed manually

in a container with a capacity of 250 L (66.04 gal.) for a period of 10 minutes. The perforated shale brick units were kept in water before they were built to lead to better bond

Mariam Farouk Ghazy is an Assistant Professor in the Faculty of Engineering in the Department of Structural Engineering at Tanta University, Tanta, Egypt. Her research interests include concrete technology, fiber-reinforced concrete, inspection and quality control of reinforced concrete, and composites.

is demanded, the incorporation of fine, flexible, and evenly distributed fibers in the lime-based mortar could be a solution.

To investigate the effects of various process parameters on the final results and then to suggest the near-optimum (the best) process settings, statistically designed experi-ments were used in this study. The Taguchi11 method, a powerful experimental design tool, uses a simple, effective, and systematic approach for setting suitable process param-eters to effectively control the amount of final results and to easily determine what parameters have the most significant effects on the final results. Further, this approach requires minimum experimental cost and efficiently reduces the effect of the source of variation. Taguchi11 has developed a system of tabulated designs (arrays) that allow for the maximum number of main effects to be estimated in an unbi-ased (orthogonal) manner, with a minimum number of runs in the experiment.

In this study, three parameters were taken with three different levels of each. Thus, a total of 27 (33) different combinations were considered according to full factorial design. According to Taguchi,11 however, the samples could be organized into only nine groups. If they were considered separately, they would still yield results with the same confidence.

Taguchi’s11 method of experimental design provides a simple, efficient, and systematic approach for the optimi-zation of experimental designs for performance quality and cost.12 The traditional experimental design methods are too complex and difficult to use. Additionally, large numbers of experiments have to be carried out. Traditional experimen-tation involves one-factor-at-a-time experiments, wherein one variable is changed while the rest are held constant. The major disadvantage of this strategy is that it fails to consider any possible interactions between the parameters. It is also impossible to study all the factors and determine their main effects—that is, the individual effects in a single experiment. The Taguchi11 technique overcomes all of these drawbacks. Compared to the conventional approach to experiments, this method drastically reduces the number of experiments that are required to model the response functions.

In this study, a new application of Taguchi’s11 method was employed to design the experimental work and determine the effect of using different types of mortar for both interface and overlays on masonry behavior. The Taguchi11 method of offline quality control has been successfully used in the design and selection of near-optimum process parameters in many areas of manufacturing processes; however, no effort

Table 1—Brick unit properties

Manufacturer No.

Dimensions, in. (mm)

Absorption,%

Compressive strength of brick, psi (MPa)

19.1 x 4.25 x 2.44(230 x 108 x 62)

9.531435.5(9.9)

29 x 4.25 x 2.56

(229 x 108 x 65)11.69

1261.5(8.7)

39.65 x 4.57 x 2.6(245 x 116 x 66)

9.411450(10)


between brick and mortar. The mortar proportions were in accordance with ECP 204-2005.14 The ratio by volume was 1:3 for cement:sand, and the water-cement ratio (w/c) was 1.3 for plain mortar while adding HRWRA for PP fiber-reinforced mortar to improve the workability. Mortar cubes with dimensions of 2.8 x 2.8 x 2.8 in. (70 x 70 x 70 mm) were made during construction of the test specimens to record the compressive strength of mortar after 28 days (refer to Table 2). For the masonry prisms, three courses of brick units were made to measure the compressive strength and seven vertical courses of brick units were made to measure the flexural bond strength after 28 days in accordance with ASTM C1314-00a19 and ASTM E518-00a,20 respectively. Figure 1 shows the specimen’s preparation and testing. Eighteen specimens were made and cured in the laboratory condition. After 2 days, masonry prisms were covered with various layers of mortars—approximately 0.4 in. (10 mm) of thickness for each layer. Masonry prism strengths were calculated using the gross area under loading. Masonry prisms were tested using a universal testing machine with a total capacity of 67.4 kips (300 kN).

Nine wall specimens 30 in. (750 mm) long, 30 in. (750 mm) high, and 4.64 in. (116 mm) wide—the width of the brick unit—were constructed using the same perfo-rated shale brick units and mortar to build brick prisms. Bed and head joint mortar had overlays approximately 0.4 in. (10 mm) thick, which were added to the walls after running bond and mortar joint pointing 2 days after construction. The walls were capped with 0.4 in. (10 mm) thick concrete beams to distribute the applied load uniformly (refer to Fig. 2). All of the walls were air-cured inside the laboratory (at a temperature of approximately 77°F [25°C] and 70% rela-tive humidity) for 28 days. A hydraulic load cell with a total capacity of 112.4 kips (500 kN) was used to test the walls after 28 days. During testing, applied loads and midheight longitudinal strain (with a gauge length of 10 in. [250 mm]) were recorded at each load stage for each specimen. The compressive strength of masonry walls was calculated by using the gross area under loading.

Plan of experimentsIn experimental investigations, the statistical design

of experiments is used quite extensively. The statistical design of experiments refers to the process planning the experiment so the appropriate data can be analyzed by the statistical method, resulting in valid and objective conclu-sions. The design of experimental methods such as facto-rial design, response surface methodology (RSM), and the Taguchi11 method are now widely used in place of the one-factor-at-a-time experimental approach, which is time-consuming and exorbitant in cost.

Design of experiment based on Taguchi’s11 technique—The major steps required for the experimental design using the Taguchi11 method are 1) establishment of the objective function; 2) identification of the factors and their levels; 3) selection of an appropriate orthogonal array (OA); 4) experi-mentation; 5) analysis of the data and determination of the near-optimum level of each factor (optimum combination); and 6) confirmation of experiment.

Taguchi11 designed certain standard OAs by which the simultaneous and independent evaluation of two or more parameters for their ability to affect the variability of a particular product or process characteristics can be done in a minimum number of tests. While there are many standard OAs available, each array is meant for a specific number of independent design variables and levels. In this study, the behavior of three control factors each at three levels was investigated. Therefore, an L9 OA was selected for this inves-tigation. The three independent variables (control factors) and their three levels are presented in Table 3. Table 4 shows the layout of the L9 OA according to Taguchi.11 A loss func-

Table 2—Characteristics of used mortars and concrete

Material Mixture proportion PP fiber,* % w/c HRWRA†, % Compressive strength, psi (MPa)

Mortar 1C:S = 1:3

(by volume)

0

1.3

0 2320 (16)

Mortar 2 1 1 2247.5 (15.5)

Mortar 3 2 1.5 2247.5 (15.5)

Concrete C:S:G = 1:1.7:3.4 (by weight) 0 0.5 1 4350 (30)*Percentage by volume of mortar. †Percentage by weight of cement.

Fig. 1—Masonry prism specimens’ preparation and testing.

Fig. 2—Masonry wall’s preparation and testing.


[ ]∆

η = η + η + η

= η − η

1 1 2 31 ,3

Effect of max min

A

A AA (3)

A control factor with the largest effect means that it has the most significant influence on the strengthening quality. ANOVA21,22 is used to discuss the relative importance of all control factors on the cutting quality and to determine which control factors have the highest effect. Parameters used in ANOVA are calculated by the following equations

(4)

29 3 2

1 1

9 2

1

1 1, , 9 3

, ,

, , 100%

m i A A mii i

T i m e T control factori

A AA A A A A

e T

S S S

S S S S S

V SV S f F

V S

= =

−=

= η = η −∑ ∑

= η − = −∑ ∑

= = σ = ×

where Sm is the average of squares of sums; SA is the sum of squares related to control factor A; ST is the sum of squares of the errors correlated to all control factors; VA is the variance related to factor A; fA is the degree of freedom for factor A; FA is the F-ratio related to control factor A; and sA is the percentage contribution related to control factor A. The values of sB and sC can be calculated similarly. The computer values for sA, sB, and sC show the relative importance of the three control factors in determining the strengthening qualities.

RESULTS AND DISCUSSIONA statistical software is used to develop the experimental

plan for the Taguchi11 method. The same software is also used to analyze the measured data. Moreover, ANOVA was used to discuss the relative importance of all control factors and their contribution.

Masonry prismTable 5 gives the experimental test results of the compres-

sive strength and flexural bond strength for the masonry prism and corresponding S/N using Eq. (1). The mean S/N of compressive strength for each level are shown in Table 6. The effect of each control factor is computed from the value of D (hmax – hmin). Based on the data presented in Table 6, the optimal compressive strength was obtained at (A2 B2 C3) 1% PP fiber-reinforced interface mortar (Level 2), 1% PP fiber-reinforced overlay mortar (Level 2), and cover with overlays of mortar in both faces (Level 3). Moreover, factor C (the number of overlay mortar) recorded a maximum value of D (2.69) and thus had the most significant influence on the compressive strength (Rank 1).

The mean S/N of the flexural bond strength for each of the three levels is shown in Table 7. The best level for each control factor is the one with the highest S/N. Based on the data presented in Table 7, factor A (type of mortar interface) had the largest effect (recorded maximum value of D 4.69) and thus had the most significant influence on the flexural bond strength (Rank 1). The optimal flexural bond strength

tion is then defined to calculate the deviations between the experimental value and the desired value. This loss func-tion is further transferred into a signal-to-noise ratio (S/N). Usually, there are three S/N available, depending on the type of characteristic; the lower-the-better (LB), the higher-the-better (HB), and the nominal-the-better (NB). In this inves-tigation, the objective was to maximize the strength and ductility; therefore, “larger-is-better” quality characteristics were selected. The logarithmic function can be calculated as follows11

21

1 110 logn

i i

SN n y=

η = = − ∑

(1)

where the quality score yi with the larger-the-better style has been assumed. The overall mean value of h over the nine experiments becomes

9

1

19 i

i=η = η∑ (2)

The effect of a control factor level is defined as the devia-tion of its related S/N h from the mean value h. For example, when the effect of Level A1 is concerned, one can note that the control factor A is at Level 1 in Experiments 1 to 3. Hence, the average hA1 and effect of A are given, respec-tively, as

Table 3—Levels of variables used in experiment

Levels(coded)

Variables (control factors)

AType of mortar

interface

BType of overlay

mortar

CNumber of overlay

mortar

1 (Control) 0% (Ordinary) 0% (Ordinary) 0 (Without overlay)

2 PP 1% PP 1% 1

3 PP 2% PP 2% 2

Table 4—L9 OA

Experiment No.

Variables (control factors)

A B C

1 (Control) 1 1 1

2 1 2 2

3 1 3 3

4 2 1 2

5 2 2 3

6 2 3 1

7 3 1 3

8 3 2 1

9 3 3 2


was obtained at 1% PP fiber-reinforced interface mortar (Level 2), 1% PP fiber-reinforced overlay mortar (Level 2), and cover with two layers of mortar (Level 3) in the front and back face. The experiment adopting the best level combi-nation was A2 B2 C3 for Experiment 5, which is listed in Table 4.

The results of the ANOVA for the masonry prisms’ compressive strength and flexural bond strength are given in Tables 8 and 9, respectively. The last column of these tables indicates the percentage of each factor’s contribution on the total variation, thus exhibiting the degree of influence on the result. Table 8 reveals that factor C (the number of overlay mortar), which reached 57.61%, made the major contribu-tion to the overall performance of masonry compressive strength. The contribution percentage for factor B (type of overlay mortar) is the smallest—6.39%—probably because, in each case, the compressive strength of these mortars is approximately the same. Moreover, Table 9 shows that factor A (type of mortar interface), which reached 64.1%, made the major contribution to the overall performance of masonry flexural bond strength. The contribution percentage for factor B (type of overlay mortar) is the smallest—2.7%.

Masonry wallsTable 10 presents the experimental test results of the

ultimate loads and the maximum longitudinal strain and calculated compressive strength for the tested masonry walls and corresponding S/N using Eq. (1). From the results, a significant strength increase was observed for all tested walls compared to control wall Experiment 1. The response table mean S/N for ultimate loads and the compressive strength are given in Tables 11 and 12,

respectively. It indicates that the S/N at each level of the control factor were changed from Level 1 to Level 3. The higher the difference, the more influential the control factor. The control factors and their interactions were sorted in rela-tion to the D values. It can be seen in Tables 11 and 12 that the strongest influence was exerted by factor C (Rank 1), factor A (Rank 2), and factor B (Rank 3), respectively. It

Table 5—Test results of masonry prism

Experiment No.

Type of interface mortar (A)

Type of overlay(B)

Numberof overlay (C)

Compressive strength, psi (MPa)

S/N for compressive strength

Flexural bond strength, psi (MPa)

S/N for flexural bond strength

1 1 1 1 377 (2.6) 8.3 55.1 (0.38) –8.4

2 1 2 2 493 (3.4) 10.63 79.75 (0.55) –5.19

3 1 3 3 548.1 (3.78) 11.55 94.25 (0.65) –3.74

4 2 1 2 594.5 (4.1) 12.26 139.2 (0.96) –0.35

5 2 2 3 703.25 (4.85) 13.71 152.25 (1.05) 0.42

6 2 3 1 449.5 (3.1) 9.83 98.6 (0.68) –3.35

7 3 1 3 565.5 (3.9) 11.82 104.4 (0.72) –2.85

8 3 2 1 507.5 (3.5) 10.88 85.55 (0.59) –4.58

9 3 3 2 609 (4.2) 12.47 91.35 (0.63) –4.01

Table 6—Response table mean S/N for compressive strength, masonry prism

Level A B C

1 10.16 10.79 9.67

2 11.93 11.74 11.78

3 11.72 11.28 12.36

hmax 11.93 11.74 12.36

hmin 10.16 10.79 9.67

D (hmax – hmin) 1.77 0.95 2.69

Rank 2 3 1

Note: Bold italic numbers denote best levels.

Table 7—Response table mean S/N for flexural bond strength, masonry prism

Level A B C

1 –5.78 –3.87 –5.45

2 –1.09 –3.12 –3.19

3 –3.82 –3.7 –2.06

D 4.69 0.75 3.39

Rank 1 3 2

Table 8—Results of ANOVA for compressive strength, masonry prism

Control factors

Degree of freedom

(f)

Sum of square (SS)

Mean of square (MS

= SS/f)

F ratio(F = MS/

MSe)

Contribution (s = SS/SST), %

A 2 0.963 0.482 3.34 27.7

B 2 0.222 0.111 0.77 6.39

C 2 2.003 1.002 6.94 57.61

Error (e) 2 0.288 0.144 — 8.3

Total 8 3.477 — — 100

Notes: MSe is error mean square; SST is total sum of square.

Table 9—Results of ANOVA for flexural bond strength, masonry prism

Control factors

Degree of freedom

(f)

Sum of square (SS)

Mean of square (MS)

Fratio

Contribution s, %

A 2 0.214 0.107 20.43 64.1

B 2 0.009 0.005 0.85 2.7

C 2 0.101 0.051 9.68 30.24

Error 2 0.01 0.005 — 2.96

Total 8 0.334 — — 100


similar trends to those of the compressive strength results of masonry prisms.

Among the different control factors, the one provided by the number of overlay mortar showed the best efficiency of strengths (factor C). A possible explanation is that the over-lays could mitigate and confine the effects of damages in the bricks and the composite could attain a higher load capacity. Thus, a sudden loss of rigidity in walls without overlays was avoided. Figure 3 shows the recorded load/longitudinal strain curves for different walls. These curves show the higher longitudinal strain values for all tested walls compared with the control wall (Experiment 1). The increase up to 10 times for Experiment 5 (2% PP fiber-reinforced mortar interface and with overlays in two faces) means a higher ductility.

was evident that factor C (number of overlay mortar) had the greatest effect on the influence of the ultimate load and compressive strength of the wall testing condition.

The ANOVA terms for ultimate loads and the compressive strength of the masonry wall are shown in Tables 13 and 14, respectively. It can be observed in these tables that factor C (type of overlay mortar) had a significant influence on the wall’s behavior. The contribution percentage is 67% for compressive strength. Factor B (type of overlay mortar) shows the smallest values at 5.72%. These results exhibited

Table 10—Test results of walls

Experiment No.

Type of interface

mortar (A)Type of

overlay (B)Number of overlay (C)

Pu, kips (kN)

S/N for Pu

Compressive strength, psi (MPa)

S/N for compressive

strength

Maximum longitudinal

strain, in. (mm)S/N for maximum longitudinal strain

1 1 1 1 40.5 (180) 45.11 304.5 (2.1) 6.44 0.0146 (0.37) –8.73

2 1 2 2 58.5 (260) 48.3 398.75 (2.75) 8.79 0.085 (2.16) 6.68

3 1 3 3 76.5 (340) 50.63 478.5 (3.3) 10.45 0.091 (2.31) 7.27

4 2 1 2 69.75 (310) 49.83 475.6 (3.28) 10.32 0.078 (1.99) 5.98

5 2 2 3 96.75 (430) 52.67 609 (4.2) 12.47 0.128 (3.25) 10.37

6 2 3 1 49.5 (220) 46.85 366.85 (2.53) 8.1 0.027 (0.69) –3.23

7 3 1 3 78.75 (350) 50.88 497.35 (3.43) 10.71 0.078 (1.99) 5.98

8 3 2 1 56.25 (250) 47.96 416.15 (2.87) 9.16 0.038 (0.96) –0.355

9 3 3 2 72 (320) 50.1 491.55 (3.39) 10.6 0.11 (2.79) 8.88

Table 11—Response table mean (S/N) for ultimate loads, masonry walls

Level A B C

1 48.01 48.60 46.64

2 49.78 49.64 49.41

3 49.65 49.19 51.39

D 1.77 1.04 4.76

Rank 2 3 1

Table 12—Response table mean (S/N) for compressive strength, masonry walls

Level A B C

1 8.56 9.156 7.89

2 10.28 10.14 9.9

3 10.16 9.71 11.21

D 1.72 0.98 3.32

Rank 2 3 1

Table 13—Results of ANOVA for ultimate loads of masonry walls

Control factors

Degree of freedom

(f)

Sum of square (SS)

Mean of square (MS) F ratio

Contribution s, %

A 2 5955.6 2977.8 3.39 12.89

B 2 1688.9 844.4 0.96 3.65

C 2 36,822.2 18,411.1 20.97 79.7

Error 2 1755.6 877.8 — 3.8

Total 8 46,222.2 — — 100

Table 14—Results of ANOVA for compressive strength of masonry walls

Control factors

Degree of freedom (f)

Sum of square (SS)

Mean of square (MS)

Fratio

Contribution s, %

A 2 0.64 0.32 3.59 21.3

B 2 0.17 0.09 0.96 5.72

C 2 2.0 1.0 11.29 67

Error 2 0.18 0.09 — 5.9

Total 8 2.99 — — 100

Fig. 3—Load versus longitudinal strain of masonry walls.


Vertical and horizontal cracks appeared on the masonry walls without overlays at the midheight of the wall and then progressed diagonally to the bottom corners. The initial crack loads were approximately 62% of the ultimate loads, whereas, in the walls with overlays, debonding of the over-lays from the brick was observed, as shown in Fig. 4, and the initial crack loads were approximately 66 to 85% of the ultimate loads. The differences of compressive strength between the masonry prisms and the walls were less than 25%.

The response table mean S/N and the results of ANOVA for maximum longitudinal strain are given in Tables 15 and 16, respectively. The optimal longitudinal strain of the control factors recorded at (A3 B2 C3) 2% PP fiber-reinforced interface mortar (Level 3), 1% PP fiber-reinforced overlay mortar (Level 2), and cover with overlays of mortar in both faces (Level 3), is given in Table 15. Table 16 reveals that factor C (number of overlay mortar), which reached 81%, made the major contribution to the overall performance; this is due to the impact of the use of PP fiber in the mortar of the overlays. The contribution percentage for factor A (type of mortar interface) is the smallest at 3.15%.

CONFIRMATION TESTSThe confirmation experiment is the final step in any

design of the experimental process. Once the optimum (most desirable) level of the design parameters was selected, the next step was to predict and verify the improvement of the quality characteristics using the optimal level of the design parameters. It is a good idea to plan on running an additional few samples at the optimum condition. These confirmation tests serve two purposes: first, they establish the new performance at the new (optimum) condition, which can establish the improvement achieved. Second, they allow the experimenter to determine how close the estimate is to the results observed. The result expected is considered to be confirmed when the mean of a number of samples tested at the optimum condition falls close to it.

The predicted (calculated) (S/N) (h^) using the optimum combination of the design parameters can be calculated as

^

1( )

n

m i mi=

η = η + η −η∑ (5)

where hm is the total mean (S/N); hi is the mean of the S/N at optimal level; and n is the number of the main factors that affects the quality characteristics. The results of the confir-mation experiment are shown in Table 17. In this table, ymean shows the arithmetic average value of y1, y2, and y3, while hver is the verification test results (S/N calculated by Eq. (1)).

Table 17 shows the comparison of the predicted values with the actual values using the optimum combinations; much lower differences were observed and the differences in all cases fall within the reasonable limit.23 The most desirable combination (the optimum combination) for the masonry prism compressive strength is determined to be at A2 B2 C3. When the process is set at this condition, it is expected to improve performance by 19.64% = (2.214/11.272) × 100, (hm = 11.272, h^ = 13.486 = 11.272 + 2.214).

As shown in this table, the experimental values agree reasonably well with the predicted values. An error of 1.27% for the S/N of compressive strength is observed when the

predicted result is compared with the experimental value of the masonry prism. Hence, the experimental result confirms the optimization of the process parameters using the Taguchi11 method for enhancing the process performance. The resulting model seems to be capable of predicting the responses of the process with reasonable accuracy.

CONCLUSIONSAccording to the experimental study in this paper, the

concluding remarks are as follows:1. The application of the Taguchi11 method for the design

of experiments is simple and efficient. Additionally, an adequate number of experiments were carried out.

2. The experimental results confirm the optimization of the process parameters using the Taguchi11 method for

Fig. 4—Testing of masonry walls.

Table 15—Response table mean S/N for maximum longitudinal strain, masonry walls

Level A B C

1 1.74 1.08 –4.11

2 4.37 5.57 7.18

3 4.83 4.31 7.87

D 3.09 4.5 11.98

Rank 3 2 1

Table 16—Results of ANOVA for maximum longitudinal strain, masonry walls

Control factors

Degree of freedom (f)

Sum of square (SS)

Mean of square (MS)

Fratio

Contribution s, %

A 2 0.23 0.11 0.51 3.15

B 2 0.72 0.36 1.64 9.82

C 2 6.12 3.1 13.92 80.99

Error 2 0.44 0.22 — 6.05

Total 8 7.52 — — 100


enhancing the process performance. The resulting model seems to be capable of predicting the responses of the process with reasonable accuracy.

3. The number of overlays and the type of interface mortar were the most influential factors on the masonry prisms’ compressive strength and flexural bond strength, respectively.

4. The application of mortar overlays increased the wall strength. The strengthening efficiency was not dependent on the overlay mortar types but instead on the number of mortar overlays.

5. PP fiber-reinforced mortars for overlays did not show remarkable efficiency in increasing the strengths of prisms and walls. The obvious efficiency in increasing the longitu-dinal strains of walls.

6. Based on the mean S/N results, the strongest influence on the compressive strength of prisms and walls was exerted by control factor C (number of overlay mortar), control factor A (type of mortar interface), and control factor B (type of overlay mortar), respectively. The optimal compressive strength recorded was at (A2 B2 C3) 1% PP fiber-reinforced interface mortar, 1% PP fiber-reinforced overlay mortar, and cover with overlays of mortar in both faces.

7. Based on results of ANOVA, the number of mortar overlays showed the best efficiency of wall strength and ductility; the contribution percentage was 67% and 81% for the compressive strength and maximum longitudinal strain of walls, respectively.

8. In a masonry wall covered on both faces with different cement mortar overlays, the overlays become part of the composite element.

9. Mortar overlays can be used to strengthen masonry walls subjected to compression loads.

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Table 17—Confirmation test results

ResponseOptimum

combination

Verification test resultsCalculated value h^

Difference |hver – h^

cal|Error,

%Improvement

ratio, %y1 y2 y3 ymean hver

Masonry prism, compressive strength

A2B2C3 4.85 4.72 4.9 4.823 13.66 13.486 0.174 1.27 19.64

Masonry walls, ultimate load (compressive strength)

A2B2C3

(A2B2C3)432

(4.31)430

(4.12)433

(4.15)431.66 (4.19)

52.7 (12.44)

52.516 (12.288)

0.184 (0.152)

0.35 (1.22)

6.9 (27)


Title no. 109-M50




Experimental Study on Dynamic Axial Tensile Mechanical Properties of Concrete and Its Componentsby Shengxing Wu, Yao Wang, Dejian Shen, and Jikai Zhou

that is, those that are always fractured by tension and show no obvious plastic deformation before fracture. Currently, dynamic axial compression experiments on concrete, rock, and mortar are relatively common, but experimental results for dynamic axial tension testing are scarce, especially for mortar and the mortar-rock ITZ due to the difficulty of axial tensile testing. Moreover, the existing studies on concrete and its three phases at the same loading conditions are very limited. In addition to being few in number, most of these studies have focused on the effects of aggregate-cement paste bond strength on the mechanical behavior of concrete and only in static compressive or flexural loads.13-15 Using aggregate strength, mortar strength, cement paste, and bond strength, Husem13 also obtained basic expressions for estimating the compressive strengths of lightweight and ordinary concretes. The results again confirmed that the components play critical roles and display regular relations determining the mechanical properties of concrete. Thus, it is very important to perform further experimental research on the dynamic axial tensile properties of concrete and its components to enable numerical analysis and further study of their relations.

Currently, there have been few numerical studies at the mesoscale investigating concrete dynamic mechanical prop-erties16,17 and, more importantly, parameters such as the dynamic strengths and moduli of the three phases in the same concrete are lacking; hence, the parameters required for the mesoscale modeling of concrete are unavailable. In this study, the authors performed systematic experiments on concrete component materials and concrete made from the same components.

Different strain rates, different initial static loads, and cyclic loads with different frequencies are common phenomena in earthquakes; these are important factors influencing the dynamic properties of quasi-brittle mate-rials. Particularly, the initial static load is usually put on a concrete structure to some degree before an earthquake happens; thus, the effects of initial static loading cannot be avoided. There have been few reports considering the initial static load, and these reports have only considered compres-sion loading.18-20 Furthermore, there has been no research on mortar and the mortar-rock interface.

A series of dynamic axial tensile tests were performed on concrete and its three components using a servo-hydraulic testing machine. The dynamic mechanical properties of approximately 200 speci-mens were tested under a dynamic load at strain rates that ranged from 10–6 to 10–2 s–1, different initial static loads, and cyclic vari-able-amplitude loads with different frequencies. The results indi-cated the following: 1) tensile strength is sensitive to strain rate in all these materials and the rate sensitivity of strength for concrete was close to the composite material with the lowest sensitivity factor k; 2) the elastic modulus is less sensitive to strain rate than strength in all the materials. The rate sensitivity of the modulus for concrete was close to its component material with the lowest sensitivity factor m. The interfacial transition zone (ITZ) had the highest m among the composites; 3) the stress-strain relation for mortar is almost completely linear before peak stress. In contrast, the stress-strain relations of the concrete, granite, and interface appear nonlinear when the stress is set at more than approxi-mately 50% of the peak value, and the nonlinear section showed a linear trend with increasing strain rate; 4) an initial static load within certain limits increased the dynamic tensile strength. The critical initial static loads for the mortar, granite, interface, and concrete were 70%, 50%, 50%, and 30%, respectively; and 5) the cyclic loading history had the least influence on the mortar and the most influence on the interface. The influence of fatigue damage decreased when the loading frequency increased.

Keywords: component materials; cyclic load; dynamic axial tension; failure mechanism; initial static load; strain rate.

INTRODUCTIONAt the mesoscale, concrete can be regarded as a three-

phase composite consisting of a coarse aggregate, a mortar matrix, and an interfacial transition zone (ITZ) between the aggregate and the mortar matrix. The macromechanical behavior of concrete greatly depends on the properties of its component materials, and research on concrete proper-ties based on mesomechanics has been quite substantial. In addition, the properties of concrete structures under dynamic loading, such as encountering an earthquake or blasting, have received continuous attention. Understanding the dynamic mechanical properties of the component materials lays the foundation for the analysis of the mechanical behavior of concrete at the mesoscopic level. Thus, further research on all the components of concrete is required.

There has been a great deal of research on concrete1-6 and rock7-10 and a small amount of research on mortar.11,12 The conclusions that the dynamic strengths of concrete and rock increase with increasing strain rate have been acknowledged by all researchers, but many systematic tests under the same loading conditions are needed to determine the relations between them and whether they share the same rate sensi-tivities to enable a complete rate effect analysis of concrete.

The three phases of concrete have similar properties as bulk concrete. They are all classified as quasi-brittle materials—


ratio of water, cement, and sand of 1:2:4. The compres-sive strength of the cubical mortar samples (with a side length of 70.7 mm [2.78 in.]) was approximately 22.8 MPa (3.32 ksi). The average density of the natural granite was 27.2 kN/m3 (0.1 lb/in.3), and the average compressive strength was 64.7 MPa (9.38 ksi). ITZ specimens were made from granite and mortar with the aforementioned propor-tions. The coarse aggregate in the concrete samples was made from the same rock as the granite matrix by crushing it. The maximum diameter of the coarse aggregate was approximately 20 mm (0.79 in.). The mass ratio of water, cement, sand, and aggregate in the concrete was 1:2:4:6—equal to the ratios in the mortar. The compressive strength of the concrete was approximately 33.5 MPa (4.86 ksi).

Treatment of specimens and testing deviceTo measure longitudinal tensile strain, two pairs of longi-

tudinal strain gauges 100 mm (3.94 in.) long were attached uniformly and symmetrically on each of the concrete, mortar, and granite sample surfaces with epoxy. As for the ITZ specimens, it was difficult to measure the actual defor-mation of the ITZ because the mortar-aggregate ITZ is less than 100 mm (3.94 × 10–3 in.) thick.21 Strain gauges 5 mm (0.197 in.) long were chosen to measure the tensile strain of the ITZ by attaching them at visible interfaces of the specimens. The strain results from this approximate measure method for the ITZ could basically meet the requirements of numerical analysis for concrete at the mesoscale because the element size of the ITZ cannot be chosen to be arbitrarily small, considering the computational limitations.

For transferring tensile loads, 40 mm (1.57 in.) thick circular steel plates were attached to the ends of the samples with an epoxy structural adhesive that has a tensile strength of 20 MPa (2.9 ksi). Then, the specimens were connected to a hydraulic clamp installed on the loading device through a spherical hinge apparatus with a screw in the center of the steel transfer plates. The tensile tests were performed on a servo-hydraulic testing system. The strains were measured with a dynamic strain instrument. Figure 2 shows the servo-hydraulic testing system and the test specimen.

Testing schemeIn this study, 10 loading conditions were applied to four

kinds of specimens, using a total of approximately 200 spec-imens. The strain rate, initial static load, and loading history were considered to be the factors influencing the dynamic tensile properties of the materials. The testing scheme is shown in Table 1. The details are as follows.

Loading at different strain rates—Monotonic loading tests were applied at strain rates of 10–6, 10–5, 10–4, 10–3, and 10–2 s–1, within the range of earthquake loading speeds. A strain rate

Shengxing Wu is a Professor and PhD Supervisor in the School of Civil Engineering at HoHai University, Nanjing, China, where he received his PhD. His research interests include durability of concrete structures and dynamic mechanical properties of concrete.

Yao Wang is an Associate Professor at Jinling Institute of Technology, Nanjing, China, and a Doctoral Student in the School of Civil Engineering at HoHai University. Her research interests include dynamic mechanical properties of concrete and test technology.

Dejian Shen is a Lecturer in the School of Civil Engineering at HoHai University. He received his PhD from Tongji University, Shanghai, China. His research interests include durability of concrete structures and dynamic mechanical properties of concrete.

Jikai Zhou is an Associate Professor in the School of Civil Engineering at HoHai University. He received his PhD from HoHai University. His research interests include dynamic mechanical properties of concrete and properties of concrete at microscale.

RESEARCH SIGNIFICANCEThe determination of the dynamic axial tensile proper-

ties of concrete and its component materials at the same loading conditions is necessary for the numerical simulation and design of concrete at the mesoscale and for the analysis of dynamic failure mechanisms in concrete. The results currently reported in the literature are scarce, however, and there has been a lack of systematic research on this topic. In this study, dynamic axial tensile tests of concrete and its corresponding components were conducted together, considering the factors of strain rate, initial static load, and cyclic load.

EXPERIMENTAL PROCEDURESpecimen fabrication

All specimens were cylindrical cores 68 mm (2.68 in.) in diameter and 160 mm (6.30 in.) tall. Rock samples were vertically cored out from a stripy granite matrix. Concrete and mortar were drilled from their casting blocks perpen-dicular to the casting direction. The core samples were all trimmed at each end to at least 20 mm (0.79 in.) beyond the laitance layer. No standard test method exists for directly measuring the mechanical properties of mortar-aggregate ITZs. Presently, researchers use indirect test methods, using “a specimen containing an ITZ” to test its macrome-chanical behavior. This method was also used in this study. For the sake of convenience, the terms “ITZ specimen” or “ITZ material” are used in this paper to describe the ITZ’s behavior. The ITZ specimens were manufactured in cylin-drical molds; the manufacturing process is shown in Fig. 1. First, granite coring was cut with a smooth surface on one end and a naturally fractured surface at the other end. Then, it was placed at one end of the mold and, finally, covered with mortar.

The mortar was made of ordinary portland cement and medium sand. All the mortar specimens had the same mass

Fig. 1—Manufacturing process of ITZ specimen.


of 10–6 s–1 is considered to be a static strain rate. The loading conditions and loading control modes of Conditions 1 to 5 are shown in Table 1, where the displacement speeds of the actu-ator are given only for general indication.

Loading with different initial static loads—The initial static load tests were at 30, 50, and 70% of the static tensile strength, and then a high-speed tensile load was applied until the specimen fractured. The loading conditions and loading control modes of Conditions 6 to 8 are shown in Table 1. Figure 3 shows the load waves with different initial static loads.

Loading with cyclic variable-amplitude loads at different frequencies—A triangular wave with variable amplitude was used to simulate an earthquake wave. A load of 50% of the static strength was monotonically applied to the specimen first at a static loading speed, and then a cyclic variable-amplitude loading was applied at frequencies of 1 or 5 Hz until sample failure. Each increase in amplitude was approx-imately 5% of the static strength. The centerline of the loading wave shows a small slope—the result of maintaining the load in the range of tension. The loading conditions and loading control modes of Conditions 9 and 10 are shown in Table 1. Figure 4 shows the load wave with cyclic variable-amplitude loading at frequencies of 1 and 5 Hz.

ANALYSIS OF EXPERIMENTAL RESULTSThe tests had a high success ratio; more than 80% of the

specimens made of concrete, mortar, and granite fractured within one-third of the middle section of the total length. Only

a few of the specimens fractured near the ends (within 2 mm [0.079 in.]); these test results were not considered. As for the ITZ specimens, most of them fractured at their interface. Only a few of them fractured at the section of mortar (more than 2 mm [0.079 in.] from the visible interface) or near the ends; these test results were also not considered. Figure 5(a) through (d) shows some of the mortar, granite, concrete, and ITZ specimens after the loading tests.

Comparing the concrete failure surfaces from the static and dynamic tests, the authors found that many mortar-aggregate ITZs in the concrete fractured when the strain rate was low, but the fracture surface became flatter when the strain rate was high and more coarse aggregate in the concrete broke, leading to a higher strength. This result is in accordance with those of most studies and has been accepted as one of the main reasons why the tensile strength of concrete increases with the strain rate.

Effect of strain rate on dynamic tensile strength and deformation

Effect of strain rate on tensile strength—Approximately 20 specimens of each kind of material were tested under Loading Conditions 1 to 5, which are listed in Table 1. The test results reflect the strain rate sensitivity of each mate-rial. Table 2 lists the average measured strain rate, the corre-sponding average tensile strength, and the dynamic increase factor (DIF) of the strength, herein defined as the ratio of dynamic strength to static strength. (The results for Condi-

Fig. 2—(a) Servo-hydraulic testing system; and (b) picture and sketch of test specimen.

Table 1—Loading conditions and loading control modes

Condition Loading condition Loading control mode

1 Quasi-static load (e· ≈ 10–6 s–1) Displacement speed of 0.02 mm/s

2 Dynamic Load 1 (e· ≈ 10–5 s–1) Displacement speed of 0.2 mm/s

3 Dynamic Load 2 (e· ≈ 10–4 s–1) Displacement speed of 2 mm/s

4 Dynamic Load 3 (e· ≈ 10–3 s–1) Displacement speed of 20 mm/s

5 Dynamic Load 4 (e· ≈ 10–2 s–1) Displacement speed of 200 mm/s (actually 70 mm/s)

6 30% initial static load + Dynamic Load 4Force rate of approximately 500 to 800 N/s for initial static load

Displacement speed of 20 mm/s for dynamic load7 50% initial static load + Dynamic Load 4

8 70% initial static load + Dynamic Load 4

9 50% initial static load + 1 Hz cyclic triangular wave Force rate of approximately 500 to 800 N/s for initial static load;force control at 1 and 5 Hz10 50% initial static load + 5 Hz cyclic triangular wave

Note: 1 mm/s = 0.03937 in./s; 1 N/s = 0.2248 lbf/s.


ranking order of strengths was almost invariant, as shown in Fig. 7(a); the order was granite > concrete > mortar > interface. The strength of the interface was merely slightly lower than that of the mortar, which contradicts the find-ings of Wong et al.22 It is thus necessary to further study the factors influencing interface strength, such as the test method, material composition, and casting quality.

3. The mortar and interface shared similar sensitivities to strain rate—approximately three times that of the rock and concrete, as shown in Fig. 7(b). In other words, the tensile strength of the mortar and interface are more sensitive to the strain rate than granite and concrete.

4. According to the test results, the strength DIF was considered to be linearly related to the natural logarithm of the strain rate increment factor. The strength DIF of all the materials can thus be expressed as Eq. (1). The test results of the fastest loading group of the mortar and the concrete were abnormal for some reasons and were excluded. The fitted lines based on Eq. (1) are shown in Fig. 6(a) through (d).

31 log( ) ( 1 10 )td d

ts s

fk

f−e

= + e < ×e�

��

(1)

where ftd is the dynamic tensile strength of the material; fts is the static tensile strength of the material at the quasi-static strain rate; ftd/fts is the strength DIF; e·d is the dynamic strain rate; e· s is the quasi-static strain rate (in this study, the lowest measured strain rate of 3 × 10–6 s–1 was defined as the static strain rate); and k is the rate sensitivity factor, which directly reflects the degree of the strain rate sensitivity of the tensile strength. By linear regression analysis, the k values of the mortar, granite, interface, and concrete were determined to be 0.3269, 0.1193, 0.3016, and 0.1044, respectively. The sequence of k was the following: interface ≈ mortar > granite ≈ concrete. The rate sensitivity of concrete was close to that of granite, which had the least sensitive factor k, in accor-dance with the characteristics of composite materials. It can be concluded that the rate sensitivity of strength for concrete was lower than the composite material.

The rate sensitivities of mortar and interface are higher than that of granite, and it seems that looser materials have higher rate sensitivity. In other words, the rate sensitivity of a material decreases with its degree of compaction. Studies at the microscale have shown that the strength of concrete is related not only to the porosity ratio but also to the pore structure.23 This observation raises the questions of whether the strain rate effect is also related to the pore structure of

tion 5 and some of Condition 4 are somewhat distorted due to the limitation of the machine’s loading speed.)

The experimental results showed the following:1. The tensile strengths of the quasi-brittle materials are

sensitive to the strain rate, showing a rising trend with the increasing strain rate. Figure 6(a) through (d) shows the vari-ation of strength with the strain rate for the mortar, granite, interface, and concrete, respectively. The standard devia-tions of the test results were less than 0.4 MPa (0.058 ksi), except in the case of the results for natural granite, which were relatively scattered.

2. Comparing the variation in tensile strength with the strain rate in all the materials, the authors found that the

Fig. 3—Loading waves with different initial static loads. (Note: 1 N = 0.225 lb.)

Fig. 4—Loading waves with cyclic loads at frequencies of 1 and 5 Hz. (Note: 1 N = 0.225 lb.)

Fig. 5—Pictures of some broken specimens.


Table 2—Average tensile strengths and strength DIFs of four materials at different strain rates

Condition No.

Mortar Granite ITZ Concrete

Measured average strain

rate, s–1

Average tensile

strength, MPa DIF


rate, s–1

Average tensile

strength, MPa DIF


rate, s–1

Average tensile

strength, MPa DIF


rate, s–1

Average tensile

strength, MPa DIF

1 3.9 × 10–6 2.31 1 4.4 × 10–6 8.87 1 5.2 × 10–6 2.03 1 7 × 10–6 3.43 1

2 4.7 × 10–5 3.40 1.47 5.9 × 10–5 9.40 1.06 2.6 × 10–5 2.39 1.18 1.4 × 10–4 3.83 1.12

3 3.9 × 10–4 3.65 1.58 6.1 × 10–4 10.60 1.21 1.6 × 10–4 2.58 1.27 7.9 × 10–4 4.16 1.21

4 2.6 × 10–3 4.49 1.94 3.9 × 10–3 11.60 1.31 9 × 10–4 3.57 1.76 2.5 × 10–3 4.33 1.26

5 4.3 × 10–3 5.44 2.35 1.2 × 10–2 12.50 1.41 6.4 × 10–3 4.03 1.99 7.8 × 10–3 5.07 1.49

Note: 1 MPa = 0.145 ksi.

Fig. 6—Variation in tensile strength with strain rate. (Note: 1 MPa = 0.145 ksi.)

Fig. 7—Tensile strength and strength DIF comparison of four materials. (Note: 1 MPa = 0.145 ksi.)

the material and whether the greater strain rate effect in mortar and interface is related to the failure located at more aggregates in concrete at higher strain rates. These ques-tions require further study of the failure mechanism. Wu et

al.24 preliminarily explored the dynamic fracture mechanism based on microstructural characteristics. Currently, there are several possible explanations for the rate sensitivity of dynamic mechanical behavior. It is generally thought that


of the strain rate, but the increase in amplitude was much lower than the tensile strength. This was because there was not sufficient time to further develop microcracks when the strain rate increased, which resulted in deformation hyster-esis and the slight increment of the deformation modulus. The results for granite were so scattered that there was no obvious rule of the rate sensitivity.

According to the test results, the elastic modulus DIF was linearly related to the natural logarithm of the strain rate, similar to the strength, as described previously. Thus, the elastic modulus DIF of all the materials can be expressed as Eq. (2). The test results of the faster loading group of the materials were abnormal for some reasons and were excluded. The fitted lines based on Eq. (2) are shown in Fig. 8(a) through (d).

30

0

1 log( ) ( 1 10 )d d

s s

Em

E−e

= + e < ×e�

��

(2)

the rate effect is related to the influence of the strain rate on crack development and can be explained from the view-point of energy dissipation and microdamage. Studies on the dynamic fracture mechanisms in these materials and their causative factors will require more experiments with new measurement techniques, such as acoustic emission25 (AE) and computed tomography (CT).26

Effect of strain rate on elastic modulus—According to the test results, the authors found that stress is almost linear with strain up to 50% of the strength. Thus, the authors took the secant modulus at 50% strength as the representative value of the elastic modulus. Table 3 lists the average elastic moduli of the four materials at different strain rates.

Figure 8(a) through (d) shows the variation of elastic modulus versus the strain rate for the concrete and its compo-nents, respectively. It was found that the elastic moduli of mortar, concrete, and the interface also increased nearly proportionally with the increment in the order of magnitude

Table 3—Average elastic modulus and modulus DIF of four materials at different strain rates

Condition No.


Measured average

strain rate, s–1

Average elastic

modulus, GPa DIF

Measured average

strain rate, s–1

Average elastic

modulus, GPa DIF

Measured average

strain rate, s–1

Average elastic

modulus, GPa DIF

Measured average

strain rate, s–1

Average elastic

modulus, GPa DIF

1 3.9 × 10–6 53 1 4.4 × 10–6 82.4 1 5.2 × 10–6 13.2 1 7 × 10–6 27.5 1

2 4.7 × 10–5 55.3 1.04 5.9 × 10–5 75 0.91 2.6 × 10–5 15.1 1.14 1.4 × 10–4 28.8 1.05

3 3.9 × 10–4 57 1.08 6.1 × 10–4 81.3 0.99 1.6 × 10–4 16.2 1.23 7.9 × 10–4 30.6 1.11

4 2.6 × 10–3 61 1.15 3.9 × 10–3 75.7 0.92 9 × 10–4 21.1 — 2.5 × 10–3 — —

5 4.3 × 10–3 — — 1.2 × 10–2 — — 0.0063 — — 7.8 × 10–3 — —

Notes: — is unmeasured items; 1 GPa = 145 ksi.

Fig. 8—Variation in elastic modulus with varying strain rate. (Note: 1 GPa = 145 ksi.)


where E0d is the dynamic elastic modulus of the material; E0s is the static elastic modulus of the material at the quasi-static strain rate; E0d/E0s is the modulus DIF; e·d and e· s are the same as in Eq. (1); and m is the modulus rate sensitivity factor, which directly reflects the degree of rate sensitivity of the elastic modulus. By linear regression analysis, the m values of the mortar, interface, and concrete were determined to be 0.0474, 0.1849, and 0.0495, respectively, approximately one-half of the rate sensitivity factors for strength. The rate sensitivity of the modulus for concrete was close to that of the composite material with the lowest m. The authors also found that the elastic modulus of concrete was lower than those of mortar and granite. It seems that the greater number of pores and microcracks contained in the interfaces hinder the transmission of rigidity.

Characteristics of stress-strain relations of materials—It was observed that there were two types of stress-strain relations for concrete and its components at different strain rates. One type is almost completely linear before the peak stress, such as that for mortar, as shown in Fig. 9(a), whereas the other type consists of two regimes: linear up to approxi-mately 50% strength and nonlinear above 50%. This two-regime relation was observed for the granite, interface, and concrete, as shown in Fig. 9(b) through (d), but the degree of nonlinearity was relatively low for concrete.

After summarizing all the results, the authors found several common characteristics of the dynamic tensile stress-strain relations in these materials:

1. Peak stress increased with increasing strain rate, whereas the strain at peak stress showed no obvious change or no obvious change regulation, except for the mortar (Fig. 9(a)), which had an increasing strain at peak stress

with the increasing strain rate. This is due to the linear stress-strain relations for the mortar, the increasing peak stress, and the increasing elastic modulus with increasing strain rate.

2. The increase in the elastic modulus with increasing strain rate was not significant.

3. For the concrete, granite, and interface, which had nonlinear stress-strain curves, there was one common feature that the nonlinear section always showed: a linear rela-tion with increasing strain rate; in other words, the tangent modulus increased, as shown in Fig. 9(b) through (d). This effect was because the influence of the strain rate on inelastic deformation was much greater than on elastic deformation, and there was not enough time to develop inelastic defor-mation, resulting in a decrease in the nonlinear deforma-tion. The failure appears more brittle, as demonstrated by the debris remaining on the fracture surfaces of specimens subjected to a high strain rate.

4. The energy absorption capacity (the area under the stress-strain curve before peak stress) of concrete and its component materials tends to increase with increasing strain rate. Table 4 lists the energy absorption capacities of the four materials at different strain rates.

Effect of initial static load on dynamic tensile properties

Effect of initial static load on tensile strength—Based on the aforementioned results, it is known that the dynamic tensile strength of these quasi-brittle materials is greater than the static tensile strength. In practical applications, however, an initial static load usually exists in the material before the dynamic load. The authors thus wondered whether this initial static load detrimentally influences the dynamic tensile strength. Approximately 12 specimens of each mate-rial were prepared for the tests in Conditions 6 to 8 (Table 1).

Fig. 9—Measured stress-strain curves of four materials at different strain rates. (Note: 1 MPa = 0.145 ksi.)


hysteresis of deformation. In addition, the cracking routing will change based on the principle of least energy consump-tion, which may cause cracks to penetrate into a higher-strength zone. On the other hand, microdamage and more microcracks appear with the increase of the initial static load, which will also result in an enhanced rate effect, whereas when the initial static load is higher than the elastic limit, nonlinear deformation develops with irreversible damage and the connection of microcracks forms a weak routing. In this case, it is the weakening effect of damage that plays an important role. Therefore, further initial static loading leads to decreased strength. The enhanced rate effect and weak-ening effect of damage for the strength were almost offset at an initial load of approximately 70%.

The critical initial static loads were almost equal to the elastic limit for the component materials of concrete, but these were lower than the elastic limit for concrete itself. This is because the initial static load of approximately 30% may have made the congenital defects (microcrack at interfaces) in concrete develop and connect together, which results in fracture along a weaker route and leads to decreasing strength.

Effects of cyclic loads with different frequencies on dynamic tensile properties

Effects of cyclic loads with different frequencies on tensile strength—Eight specimens of each material were prepared for the tests of Conditions 9 (1 Hz) and 10 (5 Hz) in Table 1.

Table 5 lists the average tensile strengths under different initial static loads for the four materials. Figure 10 shows the variation of dynamic tensile strength with the variation of initial static load for the four materials.

The test results showed that an initial static load within certain limits increased the dynamic tensile strength for the four materials, but a further increase of the initial static load caused the strength to decrease, as shown in Fig. 11. Herein, the threshold initial static load that results in a turning point of dynamic strength is called the “critical initial static load.” According to the test results, the critical initial static loads for the mortar, granite, interface, and concrete were 70%, 50%, 50%, and 30%, respectively. As for the mortar, the strength showed a rising trend with the increasing initial static load until an initial static load of 70% static strength. Tests under a larger initial static load have not been done; hence, the 70% initial static load was deemed as the critical initial static load for the mortar.

The authors also found that the dynamic tensile strength of these materials with an initial static load of 70% was almost no lower than their dynamic tensile strengths with no initial static load, except that of the ITZ (its dynamic strength was (1.76 – 1.67)/1.76 = 5.1% lower). It is suggested that a conservative engineering design, in accordance with the dynamic tensile strengths with no initial static load for these materials, can overlook the effects of initial static loads of less than 50% of the static strength.

Preliminary exploration of mechanism of initial static load effect—When an initial static load is lower than the elastic limit, the resulting deformation is reversible and small and has no obvious disadvantage on the specimens. In this circumstance, the weakening effect of damage is relatively small; it is mainly the strengthening effect of the strain rate that plays a dominant role. The authors found that the rate effect with a certain initial static load is higher than that with no initial static load, which leads to a higher dynamic strength than the pure dynamic strength. This can be explained as follows: On one hand, a sudden change in the strain rate will lead to an inertial effect, resulting in the

Table 4—Average energy absorption capacity of four materials at different strain rates (10–6 MPa)

Strain rate, s–1 Mortar Granite ITZ Concrete

10–6 50.92 1221.0 191.40 353.93

10–5 127.26 1080.63 377.47 376.44

10–4 147.12 1103.24 553.74 442.85

10–3 204.23 1314.54 — —


Table 5—Average dynamic tensile strengths of four materials under different initial static loads

Condition No.Initial static load/static

strength, %


Average strength, MPa DIF




4 0 4.49 1.94 11.57 1.31 3.57 1.76 4.33 1.26

6 30 4.78 2.07 11.50 1.30 3.72 1.83 5.00 1.46

7 50 4.78 2.07 12.51 1.41 3.69 1.82 4.63 1.35

8 70 4.95 2.14 11.45 1.29 3.40 1.67 4.39 1.27

1 100 2.31 1 8.87 1 2.03 1 3.43 1


Fig. 10—Variation of strength with different initial static loads. (Note: 1 MPa = 0.145 ksi.)


These two conditions included an initial static load of 50%. Based on the measured strain rate, the strain rates before failure of Conditions 9 and 10 for mortar were similar to and between Conditions 2 and 3, respectively, and the strain rates before failure of Conditions 9 and 10 for the other materials were similar to and between Conditions 3 and 4, respec-tively. Table 6 lists the average tensile strengths under cyclic loads of different frequencies and the relative conditions for the four materials.

The results show that the dynamic strengths in Conditions 9 and 10 were similar to and slightly lower than the strengths of the corresponding noncyclic conditions with similar strain rates. It was previously demonstrated that an initial static load of 50% is not harmful but beneficial to the dynamic strength and it was thus concluded that it is the damage from the cyclic loading that led to the decreases in strength. Comparing the four materials, the authors found that the mortar was almost unaffected by the harmful influ-ence of cyclic loading, which is in accordance with the char-acter of the linear stress-strain relation for this material. In contrast, the interface was the most affected by the cyclic loading, with the strength DIF for Condition 10 decreasing significantly (to 0.26) compared with Condition 4, which is related to the character of the obvious nonlinear stress-strain relation of the interface. The strength DIF of the other condi-tions decreased by less than 0.1.

Comparing Condition 10 with Condition 7 for the granite, interface, and concrete, the strength in Condition 10 was

Table 6—Average dynamic tensile strengths of four materials under cyclic load with different frequencies

Condition No.


Average strength, MPa DIF Average strength, MPa DIF Average strength, MPa DIF Average strength, MPa DIF

2 3.4 1.47 — — — — — —

3 3.65 1.58 10.6 1.21 2.58 1.27 4.16 1.21

4 — — 11.6 1.31 3.57 1.76 4.33 1.27

7 4.78 2.07 12.51 1.41 3.69 1.84 4.63 1.35

9 3.39 1.47 9.89 1.11 2.45 1.21 3.56 1.04

10 3.60 1.54 11.04 1.24 3.04 1.50 4.07 1.19


Fig. 11—DIF of tensile strength under different initial static loads.

much lower than for Condition 7 (the interface is still affected the most) with the same initial static load and a strain rate similar to the last cyclic test. This observation further demonstrated that low cycle fatigue damage to these quasi-brittle materials cannot be neglected.

In general, the dynamic strengths of the materials in cyclic loading were all similar to the dynamic strengths at the same strain rates, showing that dynamic strength is mainly related to the highest strain rate before failure. In addition, the dynamic strengths of the materials in cyclic loading were all lower than the strengths in monotonic loading with the same initial static load, which shows that low-frequency fatigue damage is harmful to these quasi-brittle materials. Low-frequency fatigue affected the mortar the least and affected the interface the most.

Preliminary exploration of mechanism of cyclic loading effect—Studying the stress and strain time history curves and the stress-strain curves of specimens under cyclic loading, the authors found the following:

1. There was almost no increased residual strain in the mortar with increasing cyclic loading time, which is in accordance with its linear stress-strain relation. There-fore, the influence of cyclic load for the mortar was not obvious. Figure 12(a) and (b) shows the stress-strain curves of the mortar under cyclic loading at 1 and 5 Hz; whereas the residual strain for the other three materials gradually increased with the number of cycles, this effect was more obvious at the lower frequency (1 Hz). In addition, the slope of the repeat load intends to decrease, especially in the inter-face specimens. Figure 13(a) and (b), Fig. 14(a) and (b), and Fig. 15(a) and (b) show the stress-strain curves of the granite, interface, and concrete, respectively, with 1 and 5 Hz cyclic loadings. Therefore, low-frequency fatigue damage caused a decrease in strength for these three materials and it affected the interface the most.

2. Comparing Conditions 9 and 10 for all the materials, the authors found that residual strain was reduced when the cyclic frequency was increased from 1 to 5 Hz, with the stress-strain hysteresis curve tending to close. This effect was likely because the loading and unloading speeds became much faster; thus, nonlinear deformation was reduced with increasing loading frequency, which is in accordance with the linear trend of the stress-strain curve at increasing strain rate. These results demonstrate that the harmful influence of higher-frequency cyclic loading was smaller than that of lower-frequency loading.


Fig. 12—Stress-strain curves of mortar specimens under cyclic loading. (Note: 1 MPa = 0.145 ksi.)

Fig. 13—Stress-strain curves of granite specimens under cyclic loading. (Note: 1 MPa = 0.145 ksi.)

Fig. 14—Stress-strain curves of interface specimens under cyclic loading. (Note: 1 MPa = 0.145 ksi.)

Fig. 15—Stress-strain curves of concrete specimens under cyclic loading. (Note: 1 MPa = 0.145 ksi.)


SUMMARY AND CONCLUSIONSThe dynamic mechanical properties of concrete and

its components (mortar, granite, and the mortar-granite interface) were investigated in this study. The influences of strain rate, initial static load, and cyclic loading on the tensile mechanical properties of the four materials were analyzed and compared. The main conclusions are summa-rized as follows:

1. The rate sensitivity of the dynamic tensile strength of mortar is close to that of the interface and is three times that of granite. In addition, the strength rate sensitivity of concrete is lower than that of the composites. A formula was derived to express the relationship between the strength DIF and the strain rate. The strength sensitivity factor k was proposed to compare the rate sensitivity of strength for different materials.

2. The rate sensitivities of the elastic moduli of concrete and its components are less than their rate sensitivities for dynamic strength. The rate sensitivity of modulus for the interface was the highest. The modulus rate sensitivity of concrete is close to the material that has the least sensi-tivity to the strain rate among the components. A formula was derived to express the relationship between the modulus DIF and the strain rate. The modulus sensitivity factor m was proposed to compare the rate sensitivity of the modulus for different materials.

3. The stress-strain relation for mortar is almost completely linear before peak stress. In contrast, the stress-strain rela-tions of the concrete, granite, and interface appear nonlinear when the stress is more than approximately 50% of the peak value, and the nonlinear section showed a linear trend with increasing strain rate. The degree of nonlinearity was rela-tively low for concrete.

4. An initial static load within certain limits increased the dynamic tensile strength because the enhanced rate effect played a dominant role; however, a higher initial static load reduces the dynamic strength due to the weakening effects of damage. The critical initial static loads for the mortar, granite, interface, and concrete were 70%, 50%, 50%, and 30%, respectively.

5. The dynamic strength of the materials is mainly related to the highest strain rate before failure, but low-frequency fatigue damage is harmful to these quasi-brittle materials. The influence of high-frequency cyclic loading is relatively limited. Low-frequency fatigue damage has the least influ-ence on the mortar and the most influence on the interface.

ACKNOWLEDGMENTSThe authors gratefully acknowledge the support of the National Natural

Science Foundation of China under No. 90510017 and 50979032. The authors also wish to thank the following undergraduate thesis students for their assistance during testing: Y. Wang, S. Su, W. Lu, A. Qing, Y. Chen, G. Wang, and W. Huang.

REFERENCES1. Rossi, P.; Van Mier, J. G. M.; Toutlemonde, F.; Le Maou, F.; and

Boulay, C., “Effect of Loading Rate on the Strength of Concrete Subjected to Uniaxial Tension,” Materials and Structures, V. 27, No. 5, 1994, pp. 260-264.

2. Lambert, D. E., and Ross, C. A., “Strain Rate Effects on Dynamic Fracture and Strength,” International Journal of Impact Engineering, V. 24, No. 10, 2000, pp. 985-998.

3. Brara, A.; Camborde, F.; Klepaczko, J. R.; and Mariotti, C., “Experi-mental and Numerical Study of Concrete at High Strain Rates in Tension,” Mechanics of Materials, V. 33, No. 1, 2001, pp. 33-45.

4. Georgin, J. F., and Reynouard, J. M., “Modeling of Structures Subjected to Impact: Concrete Behaviour under High Strain Rate,” Cement and Concrete Composites, V. 25, No. 1, 2003, pp. 131-143.

5. Schuler, H.; Mayrhofer, C.; and Thoma, K., “Spall Experiments for the Measurement of the Tensile Strength and Fracture Energy of Concrete at High Strain Rates,” International Journal of Impact Engineering, V. 32, No. 10, 2006, pp. 1635-1650.

6. Weerheijm, J., and Van Doormaal, J. C. A. M., “Tensile Failure of Concrete at High Loading Rates: New Test Data on Strength and Fracture Energy from Instrumented Spalling Tests,” International Journal of Impact Engineering, V. 34, No. 3, 2007, pp. 609-626.

7. Gomez, J. T.; Shukla, A.; and Sharma, A., “Static and Dynamic Behavior of Concrete and Granite in Tension with Damage,” Theoretical and Applied Fracture Mechanics, V. 36, No. 1, 2001, pp. 37-49.

8. Zhao, J., and Li, H. B., “Experimental Determination of Dynamic Tensile Properties of a Granite,” International Journal of Rock Mechanics and Mining Sciences, V. 37, No. 5, 2000, pp. 861-866.

9. Choa, S. H.; Ogatab, Y.; and Kaneko, K., “Strain-Rate Dependency of the Dynamic Tensile Strength of Rock,” International Journal of Rock Mechanics and Mining Sciences, V. 40, No. 5, 2003, pp. 763-777.

10. Asprone, D.; Cadoni, E.; Prota, A.; and Manfredi, G., “Dynamic Behavior of a Mediterranean Natural Stone under Tensile Loading,” Inter-national Journal of Rock Mechanics and Mining Sciences, V. 46, No. 3, 2009, pp. 514-520.

11. Cao, J., and Chung, D. D. L., “Defect Dynamics of Cement Mortar under Repeated Loading Studied by Electrical Resistivity Measurement,” Cement and Concrete Research, V. 32, No. 3, 2002, pp. 379-385.

12. Rome, J. I., “Experimental Characterization and Micromechanical Modeling of the Dynamic Response and Failure Modes of Concrete,” PhD dissertation, University of California, San Diego, San Diego, CA, 2002, pp. 37-77.

13. Husem, M., “The Effects of Bond Strengths between Lightweight and Ordinary Aggregate-Mortar, Aggregate-Cement Paste on the Mechan-ical Properties of Concrete,” Materials Science and Engineering, V. 363, No. 1-2, 2003, pp. 152-158.

14. Alexander, M. G.; Mindess, S.; and Diamond, S., “Properties of Paste-Rock Interfaces and Their Influence on Composite Behavior,” Materials and Structures, V. 28, No. 9, 1995, pp. 497-506.

15. Guinea, G. V.; El-Sayed, K.; Rocco, C. G.; Elices, J.; and Planas, M., “The Effect of the Bond between the Matrix and the Aggregate on the Cracking Mechanism and Fracture Parameters of Concrete,” Cement and Concrete Research, V. 32, No. 12, 2002, pp. 1961-1970.

16. Park, S. W.; Xia, Q.; and Zhou, M., “Dynamic Behaviour of Concrete at High Strain Rates and Pressures: II. Numerical Simulation,” Interna-tional Journal of Impact Engineering, V. 25, No. 9, 2001, pp. 887-910.

17. Zhou, X. Q., and Hao, H., “Mesoscale Modelling of Concrete Tensile Failure Mechanism at High Strain Rates,” Computers & Structures, V. 86, No. 21-22, 2008, pp. 2013-2026.

18. Kaplan, S. A., “Factors Affecting Relationship between Rate of Loading and Measured Compressive Strength of Concrete,” Magazine of Concrete Research, V. 32, No. 111, June 1980, pp. 79-88.

19. Yan, D., and Lin, G., “Study on Dynamic Compressive Properties of Concrete with Different Loading Paths,” Journal of Hydraulic Engineering, V. 37, No. 3, 2006, pp. 360-364.

20. Zuo, Y., “Study on Failure and Fragmentation Characteristics of Rock under Static-Dynamic Coupling Loading,” doctoral dissertation, Center South University, ChangSha, China, 2005, pp. 17-58. (in Chinese)

21. Bentur, A., “Microstructure Interfacial Effects and Micromechanics of Cementitious Composites,” Ceramic Transactions, V. 16, 1990, pp. 523-550.

22. Wong, Y. L.; Lam, L.; Poon, C. S.; and Zhou, F. P., “Properties of Fly Ash-Modified Cement Mortar-Aggregate Interfaces,” Cement and Concrete Research, V. 29, No. 12, 1999, pp. 1905-1913.

23. Zhou, J., “Mechanism and Test Study on Dynamic Flexural-Tensile Mechanic Behavior of Fully-Graded Concrete in High Arch Dam,” doctoral dissertation, HoHai University, Nanjing, China, 2007, pp. 113-125. (in Chinese)

24. Wu, S.; Zhou, J.; and Chen, H., “Unified Model and Mechanism for Dynamic Tensile Strength Increase of Concrete Base on Its Microstructure Characteristics,” Journal of Hydraulic Engineering, V. 41, No. 4, 2010, pp. 419-428. (in Chinese)

25. Wu, K.; Chen, B.; and Yao, W., “Study of the Influence of Aggre-gate Size Distribution on Mechanical Properties of Concrete by Acoustic Emission Technique,” Cement and Concrete Research, V. 31, No. 6, 2001, pp. 919-923.

26. Lawler, J. S.; Keane, D. T.; and Shah, S. P., “Measuring Three-Dimensional Damage in Concrete under Compression,” ACI Materials Journal, V. 98, No. 6, Nov.-Dec. 2001, pp. 465-475.

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Title no. 109-M51




Potential Recycling of Bottom and Fly Ashes in Acoustic Mortars and Concretesby Carlos Leiva, Luis F. Vilches, Celia Arenas, Silvia Delgado, and Constantino Fernández-Pereira

In general, to achieve an adequate environment from a noise-level point of view, noise-reducing systems providing a high level of absorption and isolation are necessary. Among these, acoustic screens or acoustic barriers are commonly used. The barriers can be composed of two different materials—namely, reflective and absorbent.

An absorbent material, such as porous concrete (PC), is a rigid acoustic-absorbing material containing large voids that have been intentionally developed for acoustic absorption or other purposes.3,4 One of the main characteristics of the acoustic absorption material commonly used outdoors is the absence or limitation of fine aggregates in the matrix. The porosity is achieved by gap-grading the coarse aggregates. The aggregates are generally joined together by means of a cement-rich mortar in a proportion of one part of cement and one or two parts of coarse aggregate, yielding a concrete showing a network of (internal) interconnected pores, also connected to the exterior, and therefore exposing the high porosity of the material.2,4,5 This is an important issue in the acoustic barriers, in which the noise-exposed surface is composed of an acoustic absorption material layer, where the incident acoustic wave penetrates within the pores, making the air vibrate; this vibration produces friction with the walls of the cavities and a loss of kinetic energy, which is transformed into heat, achieving the sound absorption.

An acoustically reflective material is rigid and acts as a good reflector of sound. Examples of these materials are wood and concrete. Thus, when sound strikes the surface of a reflective barrier, some energy is transmitted through the wall but the bulk is reflected back.6

Research on acoustic materials made of BAs and/or FAs from coal power plants is justified by the physicochemical characteristics of these by-products and may have great interest in the field of using recycling materials in acoustic barriers. In general, BA and FA from coal power plants are constituted by mixtures of oxides of various elements with a very small unburned matter content, which may be catalogued as inert waste7 and are usually destined to be landfilled. The acoustic and non-acoustic specifications demanded of acoustic barriers8-10 could, in principle, be met or even improved by the use of BA and/or FA.

To better understand the recycling potential of the ashes in this field and as an example, taking into account the height and density of acoustic barriers (screens) currently used to avoid traffic noise, one can calculate that the total annual

This study performed an evaluation of the physical, mechanical, and sound absorption characteristics of mortars and concrete containing co-combustion bottom (BA) and fly ashes (FA). The objectives were to produce building elements capable of reducing noise and to use co-combustion residues. The obtained results demonstrated that the use of BA in a concrete formula-tion (up to 60 wt%) produced an increase of sound absorption capacity—similar to the sound absorption capacity observed in porous concrete (PC) commonly used in acoustic barriers—but the mechanical strength decreased. The FA mortars presented a high reflection coefficient (RFC) and also showed reduced mechanical strength, similar to the BA concrete. According to different leaching test results, no problems were found in this product.

Keywords: bottom ash; fly ash; leaching; reflection coefficient; sound absorption.

INTRODUCTIONThe protection of the environment should be promoted by

the recovery of waste materials and their use as secondary raw materials. In many cases, recycled materials must compete with low-cost materials. When the properties of waste products make their use possible in high-added-value applications, however, these products can success-fully compete with products made from primary materials, reducing the environmental costs of waste disposal.

Pulverized coal firing is by far the most common form of coal combustion used today for power generation. The bulk of the ash (approximately 90%) that is formed when pulver-ized coal burns is converted into fine dry powder usually known as fly ash (FA) or pulverized FA. A smaller propor-tion (approximately 10%) of the coal ash exits the boiler as furnace bottom ash (BA) or slag. It is much denser and larger in size. Effectively, it is fused ash that has agglomerated and fallen down to the base of the furnace.1

With the advent of pulverized coal firing systems, signifi-cant quantities of combustion residues became available. The amounts have increased dramatically during the last few decades; for example, the quantity of ash and slag produced in the European Union (EU-15) in 2007 was 41.8 million tonnes (FA, 9.21 × 1010 lb) and 5.7 million tonnes (BA, 1.25 × 1010 lb) according to the European Coal Combustion Prod-ucts Association (ECOBA) (www.ecoba.com). It is evident that simply discarding such quantities of material is unac-ceptable on environmental grounds, so finding methods for their reuse are welcome and very necessary.

Among the environmental problems affecting quality of life, noise due to traffic (cars, trains, and planes) must be emphasized. To overcome this problem, new and better materials need to be developed to produce a greater reduc-tion in noise2; however, solutions that involve applying materials to reflect noise instead of using materials to absorb it only minimize the problem.


In addition, ordinary portland cement Type II (CPII) (CEM II/B-L 32.5N according to EN 197-111), fine aggregate (Fine), and coarse aggregate (Coarse), in the form of natural river sand and crushed granite, respectively, were used.

The chemical composition, in accordance with ASTM D3682-0112 for the different materials, is shown in Table 1.

As also shown in Table 1, the sum of the percentages of SiO2, Al2O3, and Fe2O3 reach 80.89% in the FA, indicating that it can be classified as an F-type ash, as prescribed by ASTM C618-05.13 The calcium content of the BA is low (<10%) and the sum (SiO2 + Al2O3 + Fe2O3) reaches 86.7%. Based on a chemical equivalency, this BA could meet the ASTM C618 requirements for an F-type ash. The fine and coarse aggregates are fundamentally composed of SiO2; all other components analyzed remain insignificant. Figure 1 shows cumulative percentage non-retention curves for five different materials in a semi-logarithmic scale. It can be seen that FA and CPII are the finest materials and BA presents a size distribution intermediate between fine and coarse aggregates.

A constant small amount of exfoliated vermiculite (V) was added to the mixture. V is commonly used as an additive in sound absorption materials14 and increases the mechan-ical properties of the construction materials with BAs and FAs.15 V is a hydrated silicate comprising magnesium, aluminium, and iron, which has a flaky structure. V is also usually added to mortars used for fire and acoustic protec-tion, as indicated in previous papers by the authors.16,17 The V used in this study is commercial V with 84.9% of particles less than 1.41 mm (4.6 × 10–3 ft) in size.

Preparation of test specimensThe authors’ goal was to analyze the influence of the FA

and BA studied in the properties of the acoustic absorption material prepared therewith. The different concrete and mortar compositions are shown in Table 2. In all the ash mortar and concrete compositions, the V content was kept constant and equal to 20 wt%.

To compare the properties of these materials with other concrete products usually employed in acoustic barriers, a standard concrete (SC) and a PC were also manufactured. The composition of these concretes is shown in Table 3. The proportions of CPII and coarse aggregate in PC were

Carlos Leiva is an Assistant Professor at the University of Seville, Seville, Spain. He received his BS and MS in industrial engineering and his PhD in chemical engineering from the University of Seville in 2001, 2003, and 2006, respectively. His research interests include recycling of waste in construction materials (fire-resistant and acoustic absorption materials).

Luis F. Vilches is an Associate Professor at the University of Seville. He received his BS and MS in industrial engineering and his PhD in industrial engineering from the University of Seville in 1990, 1998, and 2002, respectively. His research interests include recycling of industrial by-products and water and wastewater treatment.

Celia Arenas is a Chemical Engineer at the University of Seville. She received her BS and MS in chemical engineering from the University of Seville in 2008 and 2010, respectively. Her research interests include recycling of waste in construction materials (fire-resistant and acoustic absorption materials).

Silvia Delgado is a Chemical Engineer at the University of Seville. She received her BS in chemical engineering from the University of Seville in 2009. Her research inter-ests include recycling of waste in different applications.

Constantino Fernández-Pereira is a Full Professor in the Department of Chemical and Environmental Engineering at the University of Seville. He received degrees in chemical sciences and environmental engineering and his PhD in chemistry in 1983 from the University of Seville. His research interests, in addition to educational issues, include industrial solid waste engineering: waste characterization, waste treatment (also including wastewater treatment), and waste recycling and valorization.

BA production in a 550 MWe pulverized coal power plant could be recycled in approximately 60 km (196,850 ft) of road using concrete acoustic barriers, whenever the use of a concrete containing 50 to 60% of slag was justified because it met the technical specifications required.

RESEARCH SIGNIFICANCEThe main objective of this study is to present a series of

physical, mechanical, acoustic, and environmental proper-ties of products manufactured using different contents of FAs and BAs, with the aim of analyzing the influence of by-products that could be recycled as acoustic barriers. This study may be considered a first step in the technological development of new materials with a potential application in the field of acoustic protection against noise.

EXPERIMENTAL PROCEDUREMaterials

In this study, FA and BA from the co-combustion of coal and petroleum coke (70/30) in a power plant were studied.

Table 1—Chemical composition (wt%)

FA BA CPII Fine Coarse

SiO2 48.72 52.32 13.83 96.21 85.73

Al2O3 24.26 25.14 3.53 0.76 4.96

Fe2O3 7.91 9.23 2.26 0.22 2.92

MnO 0.07 0.07 0.06 <0.01 0.04

MgO 1.78 1.84 0.7 0.01 0.30

CaO 2.26 2.37 59.33 0.13 0.46

Na2O 0.71 0.66 0.08 0.05 1.14

K2O 3.69 3.72 0.48 0.30 0.99

TiO2 1.51 1.45 0.19 0.12 0.23

P2O5 0.35 0.25 0.06 0.01 0.06

SO3 0.02 0.03 1.68 0.02 0.03

Loss on ignition (LOI) 6.6 1.1 15.5 0.3 0.9

Specific gravity 2.7 2.3 3.1 2.8 2.6


chosen, optimizing the sound absorption coefficient (SAC) of different mixtures.

The solid components shown in the tables mentioned previously were placed in a concrete mixer and mixed until a homogeneous dry mixture was achieved. Then, water was added to the mixture and was mixed again until a homo-geneous wet mixture was obtained. In all cases, the water-solids ratio (w/s) was kept constant at 0.4. When the mixing was completed, the mixture was placed in molds and it was consolidated twice by vibration using a vibration table at half-full and again when full.

The samples were taken out of the molds after 24 hours and were cured at ambient temperature for more than 28 days (average temperature: 20°C (68°F); average relative humidity: 45%). The cured samples were used to make test pieces of different shapes and sizes, which were used in the acoustic, physical, and mechanical tests.

Test methodsAcoustic properties—When a sound wave strikes a material

(incident energy), a portion of the sound energy is reflected back (reflected energy), a portion is absorbed by the material (absorbed energy), and a portion is transmitted through the material. The absorption coefficient is the ratio of the absorbed energy to the total incident energy. The reflection coefficient (RFC) is the ratio of the reflected energy to the total incident energy.

To determine the acoustic properties of the products prepared, the SAC and RFC were determined by the imped-ance tube method18 in samples of 40 mm (0.13 ft) thick. For the high acoustic frequency range between 800 and 5000 Hz, a 30 mm (0.09 ft) diameter tube was used; and for the medium and low acoustic frequency range between 125 and 2000 Hz, a 60 mm (0.19 ft) diameter tube was adopted. The circum-ferential edge of the test sample was carefully sealed with petroleum jelly, as recommended in EN ISO 10534-2,18 to ensure a good fit between the sample and the tube. A prelimi-nary study was carried out in an impedance tube to analyze the influence of the fit as in other previous studies.19 Each value represents the average value obtained from testing three samples of each mixture design.

Physical and mechanical propertiesWith the aim of characterizing the physical and mechan-

ical properties of the product, the following tests were carried out.

The density r of the mortar was measured by weight and volume (dimensions) measurements. Density was deter-mined in the same specimens used for acoustic properties before the acoustic test.

The method of vacuum saturation, as described in RILEM CPC 11.3,20 was followed in the determination of the open void ratio (VR) (in %). The samples were dried in an oven at 105°C ± 5°C (221°F ± 41°F) until no change in measured weight (W1) was noticed. The specimens were then placed dry in a vacuum chamber for 3 hours under vacuum, followed by total submersion in water for an additional 6 hours under vacuum, followed by continued total submersion in water for 18 hours without vacuum. The saturated surface-dry weight (W2) was then determined. The void ratio was calcu-lated using Eq. (1).

2 1

1(%) 1 100

W WVR

V −

= − × (1)

where V1 is the volume of the specimen. Three specimens of each type were used.

The compressive (RC) and bending (RF) strengths of the samples were also performed based on ASTM C39/C39M-05e221 and ASTM C348-08,22 respectively, using a compressing test machine. Three specimens of each type were tested.

Environmental studyTo facilitate their use as construction materials, the prod-

ucts developed in this study must guarantee a low toxicity level, which is often assessed through a leachability study. The environmental study was carried out using the EN 12457-4 leaching test23 to characterize the FA and BA metal leachability and evaluate the possible applications of the composite products manufactured therewith. The PC

Fig. 1—Solids grain size distribution, %. (Note: 1 µm = 0.000039 in.)

Table 2—Mixture (wt%) proportion of new materials

Mixture CPII V FA BA

FA-20 60 20 20 —

FA-40 40 20 40 —

FA-60 20 20 60 —

BA-20 60 20 — 20

BA-40 40 20 — 40

BA-50 30 20 — 50

BA-60 20 20 — 60

Table 3—Mixture (wt%) proportion of reference materials

CPII Fine Coarse

SC 20 50 30

PC 20 — 80


column test was performed in which seven eluate frac-tions are collected within the range of L/S (liquid/solid) = 0.1 to 10 L/kg (0.0016 to 0.16 ft3/lb). The total test duration is 21 days. The leachant is a preconditioned water at pH = 4. The test material is applied as received and the upflow (14 mL/h [0.0005 ft3/h]) is applied through a column waste height of 28 cm (0.92 ft) and a diameter of 10 cm (0.33 ft).

Metal analysis in leachates was carried out using atomic absorption spectrophotometry and inductively coupled plasma techniques in the Microanalysis Service of Seville University (CITIUS).

EXPERIMENTAL RESULTS AND DISCUSSIONFA mortars—influence of FA content

Table 4 shows the open void ratio and density of some of the products containing different FA proportions. As can be seen, the void ratio in all the cases is very similar due to the similar size distribution of FA and CPII. The density decreased slightly when the FA content was increased due to the lower specific density (refer to Table 1) of the FA as compared with the CPII density. The densities of the FA mortars are lower than those of the SCs and PCs due to the V effect.

Regarding the mechanical properties studied, Table 4 shows the main results and it can be seen that the compres-sive strength at 28 days decreases when the FA content is increased. The mechanical properties of the ash mortars are much lower than those measured in SC, but similar to those manifested by the PC.

Figure 2 shows the results of the SAC measurements of the FA mortars series. As can be seen in this figure, the absorp-tion coefficient is almost the same in all the cases and is very similar to that shown by SC, although it is very low when a comparison is made with the SAC of PC. Despite the fact that the density of the FA mortars is lower than that of PC, these mortars exhibit a lower SAC due to the lower porosity.

Figure 3 shows the RFC of the FA mortars. Regarding this property, it can be observed that these composite materials present a high RFC similar to that observed in SC, a material commonly used in acoustic barriers as a reflective plate.25

BA concrete—influence of BA contentTable 5 shows the open void ratio, density, and mechanical

properties of some concrete products containing a different amount of BA. When the BA content is increased, the VR tends to increase due to the particle size of the BA compared to portland cement. Comparing the values obtained in BA concrete with Specimen PC, the VRs of the samples approaching 60% of BA have similar values to those of PC.

As can be seen, the density decreases when the BA content is increased due to higher porosity and its lower specific density (Table 1). The density of the BA products is lower than that of SCs and PCs and is similar to those found in FA mortars.

As can be seen in Table 5, the compressive and flexural strengths of the BA products are reduced appreciably when the BA content is increased, and they are much lower than the mechanical properties of the SCs and PCs. Under the load, the presence of voids acts as a weakness in the cement matrix, which creates localized areas of stress and propaga-tion of crack formation.26,27 BA contents higher than 50 wt% produce materials with inadequate mechanical properties for use in this kind of application because lower mechanical properties may produce problems in the installation and,

used, a commercialized product, was also subjected to the same test to compare the leaching results.

In the Netherlands, the leaching behavior obtained by the Dutch Column24 test is a decisive method of determining if and in what way the by-products can be used as a building material or in what way the waste has to be dumped. A

Table 4—Influence of FA content on physical and mechanical properties

VR, % r, kg/m3 (lb/ft3) RC, MPa (ksi) RF, MPa (ksi)

FA-20 10.3 946 (59) 4.7 (0.68) 3.3 (0.47)

FA-40 10.1 905 (56) 3.9 (0.56) 2.6 (0.38)

FA-60 10.1 824 (51) 2.1 (0.30) 1.4 (0.20)

SC 9.2 2105 (131) 23.1 (3.35) 4.3 (0.62)

PC 26.2 1858 (115) 6.3 (0.91) 1.6 (0.23)


Fig. 2—Influence of FA content on SAC.

Fig. 3—Influence of FA content on sound RFC.

Table 5—Influence of BA content on physical and mechanical properties

VR, % r, kg/m3 (lb/ft3) RC, MPa (ksi) RF, MPa (ksi)

BA-20 11.3 912 (57) 3.4 (0.49) 3.2 (0.47)

BA-40 13.5 873 (54) 1.4 (0.20) 1.3 (0.19)

BA-50 17.1 745 (46) 1.0 (0.14) 0.9 (0.13)

BA-60 21.1 701 (43) 0.4 (0.06) 0.4 (0.06)

SC 9.2 2105 (131) 23.1 (3.35) 4.3 (0.62)

PC 26.2 1858 (115) 6.3 (0.91) 1.6 (0.23)



during the service life of the acoustic-insulating element in relation to its impact resistance, freezing-and-thawing dura-bility and mechanical resistance in general.9

Figure 4 shows the results obtained after the sound absorp-tion study carried out on the BA products. The BA produced an increase of the SAC at all the frequencies compared to those of FA products due to the high porosity developed in the material (Table 5). Inside the pores, the kinetic energy of the sound wave is transformed into heat due to the wave interaction with the walls of the chambers of the absor-bent material,2,3 so it can be expected that an enhanced porous material results in a better absorbent material. The BA product coefficients are even higher than those of SC and similar to those found in other composite materials containing wastes, such as cork or rice straw-wood, which are used in similar building applications.28,29 However, the SAC values of the BA products are slightly lower than those of the PCs, especially at low and medium frequencies. It has been stated that for PC to be effective, 15 to 25% of the open void ratio is needed. The open void ratio is the most impor-tant key factor (together with the thickness of the porous layer) that dictates the efficiency of the porous material.30 Due to this fact, when the VR is increased, the SAC values for each frequency also increase for all the compositions, as can be seen in Table 5 and Fig. 4.

Figure 5 shows the RFCs of the BA concrete; as can be seen in the figure, the reflection decreased when the BA content increased.

Environmental studyAs the Introduction to EN 1794-210 indicates, “While

performing their primary function, road traffic noise reducing devices should not pose hazards to road users or other people in the vicinity or to the environment at large.” EN 1794-2 establishes that road traffic noise-reducing

devices must specify any physical or chemical condition that could cause environmental problems. This regulation also mentions: “They (the noise-reducing devices) should be made from materials which do not emit noxious fumes or leachates as the result of natural or industrial processes or as the result of fire.” In the cases of FA and BA (and, in general, when recycled materials are used), the main envi-ronmental problem could be the release of heavy metals into the environment through leaching. When the leachability was assessed, however, neither standard leaching tests nor any pollutant concentration limits restricting their use in this kind of application could be found.

Therefore, the BA and FA were submitted to the EN 12457-423 test—a leaching test commonly used in the waste management field. Moreover, the same test was applied to the other individual components used in this study, such as CPII, Fine, and Coarse, to compare the leaching results of recycled materials with those obtained in the case of the virgin raw materials commonly used in this field. These results serve to establish a maximum leaching limit because when they are mixed with cement, a phenomenon of stabi-lization/solidification of BA and FA containing metals is produced, thus reducing the leaching of heavy metal leach-ates from the final material.31-33 The leachate results obtained are shown in Table 6.

As shown in Table 6, the metal concentration results of the BA leachate are similar or even lower, except V, than those observed for building materials traditionally used in the manufacturing of sound insulation elements. With regard to FA, it must be noted that its EN 12457-423 leachate shows relatively high concentrations of As, Cr, Sb, Ba, and espe-cially Mo. Notwithstanding, it must be emphasized that the recycling of this by-product is widely employed, principally in the manufacture of ordinary portland cement Type II, as indicated in EN 450-1,34 which authorizes the use of FAs in

Fig. 4—Influence of BA content on SAC.

Table 6—EN 12457-423 leachability of FA and BA: comparison with typical construction materials (ppb)

As Cd Cr Cu Hg Mo Ni Pb Se Zn Ba V Sb

FA 21 ≤3 170 ≤1 <30 976 ≤10 ≤3 ≤25 ≤1 317 ≤20 114

BA ≤10 ≤3 ≤1 ≤1 ≤30 ≤10 ≤10 ≤3 ≤25 ≤1 3.7 94 ≤20

CPII ≤10 ≤3 436 ≤1 ≤30 ≤10 ≤10 ≤3 ≤25 105 313 ≤20 ≤20

Fine ≤10 ≤3 ≤1 ≤1 ≤30 ≤10 ≤10 ≤3 ≤25 670 82.2 ≤20 ≤20

Coarse ≤10 ≤3 ≤1 ≤1 ≤30 ≤10 ≤10 ≤3 ≤25 18 23.4 ≤20 ≤20

Fig. 5—Influence of BA content on sound RFC.


3. The leachability study carried out to assess the envi-ronmental impact derived from the use of the by-products showed that their metal leachability is similar to that found in other traditional materials used in construction, although the FA studied exhibited high concentrations of Cr, Ba, Sb, and Mo compared with other typical construction materials used in acoustic barriers. Conversely, BA presents no signifi-cant environmental problems.

These results constitute an opportunity for the manufac-ture of an acoustic absorption product composed partly of BA from power plants, optimizing the product’s composition in light of the results obtained in this study, with sound absorp-tion capacity similar to that found in PC and acceptable mechanical properties without any relevant environmental problems. Finally, the problem of PC structures is their durability during their service life; the use of BA improves some durability properties of SC,34,35 but its resistance to chemical and physical attacks, which lead to a deterioration and variation of the acoustic properties of concrete, should be analyzed.

ACKNOWLEDGMENTSThe authors acknowledge the financial support of this research by the

Spanish Ministry of Science and Technology with European Fund for Economic and Regional Development (FEDER) funds under the Recycling of Bottom and Fly Ashes from Several Thermal Processes in Noise Reduc-tion Devices (RUIDRES) Project (CTM2007-62031).

REFERENCES1. Thompson, A., Combustion Residues: Current, Novel and Renewable

Applications. Chapter 1: Current and Future Nature of Combustion Ashes, John Wiley & Sons, Inc., Chichester, UK, 2008, pp. 1-84.

2. Park, S. B.; Seuk, D. S.; and Lee, J., “Studies on the Sound Absorp-tion Characteristics of Porous Concrete Based on the Content of Recycled Aggregate and Target Void Ratio,” Cement and Concrete Research, V. 35, No. 9, 2009, pp. 1846-1854.

3. Kim, H. K., and Lee, H. K., “Acoustic Absorption Modeling of Porous Concrete Considering the Gradation and Shape of Aggregates and Void Ratio,” Journal of Sound and Vibration, V. 32, No. 7, 2010, pp. 866-879.

4. Neithalath, N.; Marolf, A.; Weiss, J.; and Olek, J., “Modeling the Influ-ence of Pore Structure on the Acoustic Absorption of Enhanced Porosity Concrete,” Journal of Advanced Concrete Technology, V. 3, No. 1, 2005, pp. 29-40.

5. Brennan, M. J., and To, W. M., “Acoustic Properties of Rigid-Frame Porous Materials—An Engineering Perspective,” Applied Acoustics, V. 62, No. 7, 2001, pp. 793-811.

6. Butler, G. F., “Improvements in/or Relating to Noise Barriers,” Patent No. 1451193, Application 46350/72, 1972, 24 pp.

7. EULFD, Council Directive 1999/31/EC of 19 December 2002 estab-lishing criteria and procedures for the acceptance of waste at landfills pursuant to Article 16 of and Annex II to Directive 1999/31/EC.

8. EN 14388, “Road Traffic Noise Reducing Devices: Specifications,” European Committee for Standardization, Brussels, Belgium, 2006, 40 pp.

9. EN 1794-1, “Road Traffic Noise Reducing Devices—Non-Acoustic Performance. Part 1: Mechanical Performance and Stability Requirements,” European Committee for Standardization, Brussels, Belgium, 2003, 24 pp.

10. EN 1794-2, “Road Traffic Noise Reducing Devices—Non-Acoustic Performance. Part 2: General Safety and Environment Requirements,” European Committee for Standardization, Brussels, Belgium, 2003, 25 pp.

11. EN 197-1, “Cement—Part 1: Composition, Specifications and Conformity Criteria for Common Cements,” European Committee for Stan-dardization, Brussels, Belgium, 2001, 14 pp.

12. ASTM D3682-01, “Standard Test Method for Major and Minor Elements in Combustion Residues from Coal Utilization Processes,” ASTM International, West Conshohocken, PA, 2001, 7 pp.

13. ASTM C618-05, “Standard Specification for Coal Fly Ash and Raw or Calcined Natural Pozzolan for Use in Concrete,” ASTM International, West Conshohocken, PA, 2005, 3 pp.

14. Qiao, D.; Wei, J.; and Wang, L., “Effect of Vermiculite on Properties of Rubber Sound Insulator,” Reguxing Shuzhi, V. 19, No. 6, 2004, pp. 22-23.

15. Vilches, L. F., “Desarrollo y método de evaluación de nuevos productos aislantes e ignífugos obtenidos a partir de materiales residuales,” doctoral thesis, University of Seville, Seville, Spain, 2002, 272 pp.

cement and concrete but does not contemplate any leachate limits. The high Cr leachability in CPII must also be stressed, contrasting with the low Cr release observed in the case of the rest of the materials analyzed.

Table 7 shows the results of the column leachability test and the comparison with the limits established in the Soil Quality Decree35; this establishes two different categories: Category I (unrestricted use) and Category II (with restricted use if isolation measures are taken). BA is characterized by a low overall solubility and very low releases of elements of environmental concern and the column test results did not exceed any limits for Category I. FAs exhibited Cr, Mo, Se, V, and SO4

2– values in the ranges above the inert category (Category I). In any case, the Category II limits are not exceeded for these FAs and BAs.

CONCLUSIONSThis paper discusses the effect on the acoustic, physical,

and mechanical properties of mortar or concrete prod-ucts containing co-combustion by-products (BA and FA), which can be used as acoustic barriers or other sound-insulating elements.

1. No appreciable differences were found in the SACs and density between different FA contents. The sound absorption of the FA mortars is low, although such products need much improvement to behave as an acceptable noise-absorbing material; FA mortars can be used as reflection material in acoustic barriers. The mechanical properties of the mortars significantly diminished when cement was replaced by FA.

2. The BA produces an increase in the SACs of the concretes for all the frequencies, probably due to the increase of material porosity, manifested by its low specific gravity and high porosity. The mechanical properties also diminish when the BA content is increased. BA contents greater than 50% produce a very high-strength reduction and a material with a minimum load-bearing (supporting) capacity.

Table 7—NEN 734524 leachability of FA and BA: comparison with soil quality decree limits (ppm)

Elements Category I Category II FA BA

As 0.8 7.0 <0.17 <0.01

Ba 6.5 166 <0.04 <0.85

Cd 0.02 0.06 <0.01 <0.01

Cr (total) 0.36 12 1.69 <0.02

Cu 0.33 3.3 <0.02 <0.02

Hg 0.02 0.08 <0.01 <0.01

Mo 0.51 2.5 2.14 0.11

Ni 0.70 3.5 0.13 <0.05

Pb 0.99 8.1 <0.02 <0.02

Sb 0.09 1.2 <0.07 <0.03

Se 0.10 0.28 0.16 <0.05

Zn 2.3 14 0.15 <0.05

V 3.2 96 1.64 2.87

Ba 6.5 166 0.85 0.03

F– 11.4 288 21.42 <2.0

Cl– 561 8795 1.72 <1.0

SO42– 3282 66,022 3918.48 68.37


16. Leiva, C.; Vilches, L. F.; Vale, J.; and Fernández-Pereira, C., “Influ-ence of the Type of Ash on the Fire Resistance Characteristics of Ash-Enriched Mortars,” Fuel, V. 84, No. 11, 2005, pp. 1433-1439.

17. Vilches, L. F.; Leiva, C.; Olivares, J.; Vale, J.; and Fernández-Pereira, C., “Passive Fire Protection of Metal Sections by Means of a Sprayed Coal Fly Ash Mortar,” Materiales de Construcción, V. 55, 2005, pp. 25-37.

18. EN ISO 10534-2, “Acoustics Determination of Sound Absorption Coefficient and Impedance or Admittance by the Impedance Tube—Part II: Transfer Function Method,” European Committee for Standardization, Brussels, Belgium, 1998, 28 pp.

19. Kim, H. K., and Lee, H. K., “Influence of Cement Flow and Aggregate Type on the Mechanical and Acoustic Characteristics of Porous Concrete,” Applied Acoustics, V. 71, No. 7, 2010, pp. 607-615.

20. RILEM CPC 11.3, “Absorption of Water by Immersion under Vacuum,” Materials and Structures, V. 17, 1984, pp. 391-394.

21. ASTM C39/C39M-05e2, “Standard Test Method for Compressive Strength of Cylindrical Concrete Specimens,” ASTM International, West Conshohocken, PA, 2005, 7 pp.

22. ASTM C348-08, “Standard Test Method for Flexural Strength of Hydraulic-Cement Mortars,” ASTM International, West Conshohocken, PA, 2008, 6 pp.

23. EN 12457-4, “Characterization of Waste: Leaching—Compliance Test for Leaching of Granular Waste Material and Sludge,” European Committee for Standardization, Brussels, Belgium, 2003, 33 pp.

24. NEN 7345, “Leaching Characteristics of Solid Earthy and Stony Building and Waste Materials; Leaching Test; Determination of the Leaching of Inorganic Components from Granular Materials with the Column Test,” Netherlands Standardization Institute, Delft, the Netherlands, 1995, 29 pp.

25. Tomis, V., “Fully Absorptive Acoustic Barrier,” Patent No. WO 2006/081778 A1, 2006, 20 pp.

26. Goodier, J. N., “Concentration of Stress around Spherical and Cylin-drical Inclusions and Flaws,” Journal of Applied Mechanics, V. 55, No. 7, 1933, pp. 39-44.

27. Ryshkewitch, E., “Compression Strength of Porous Sintered Alumina and Zirconia,” Journal of the American Ceramic Society, V. 36, No. 2, 1953, pp. 65-68.

28. Hernández-Olivares, F.; Bollati, M. R.; Del Rio, M.; and Parga-Landa, B., “Development of Cork-Gypsum Composites for Building Applications,” Construction & Building Materials, V. 13, No. 4, 1999, pp. 179-186.

29. Yang, H.; Kim, D.; and Kim, H., “Rice Straw-Wood Composite for Sound Absorbing Wooden Construction Materials,” Bioresource Tech-nology, V. 86, No. 2, 2003, pp. 117-121.

30. Marolf, A.; Neithalath, N.; Sell, E.; Wegner, K.; Weiss, J.; and Olek, J., “Influence of Aggregate Size and Gradation on Acoustic Absorption of Enhanced Porosity Concrete,” ACI Materials Journal, V. 101, No. 1, Jan.-Feb. 2004, pp. 82-91.

31. Vilches, L. F.; Leiva, C.; Vale, J.; Olivares, J.; and Fernández-Pereira, C., “Fire Resistance Characteristics of Plates Containing a High Biomass-Ash Proportion,” Industrial & Engineering Chemistry Research, V. 46, No. 14, 2007, pp. 4824-4829.

32. Leiva, C.; Garcia Arenas, C.; Vilches, L. F.; Vale, J.; Gimenez, A.; Ballesteros, J. C.; and Fernández-Pereira, C., “Use of FGD Gypsum in Fire Resistant Panels,” Waste Management, V. 30, No. 6, 2010, pp. 1123-1129.

33. Luna, Y.; Fernández Pereira, C.; and Vale, J., “Stabilization/Solidi-fication of a Municipal Solid Waste Incineration Residue Using Fly Ash-Based Geopolymers,” Journal of Hazardous Materials, V. 185, No. 1, 2011, pp. 373-381.

34. EN 450-1, “Fly Ash for Concrete—Part 1: Definitions, Specifications and Conformity Criteria,” European Committee for Standardization, Brus-sels, Belgium, 2006, 32 pp.

35. “Decree on Soil Quality,” Staatsblad 2007, houdende regels inzake de kwaliteit van de bodem (Besluit bodemkwaliteit), Staatsblad, 2007, 179 pp.


Title no. 109-M52


ACI Materials Journal, V. 109, No. 5, September-October 2012.MS No. M-2011-196 received June 27, 2011, and reviewed under Institute publication

policies. Copyright © 2012, American Concrete Institute. All rights reserved, including the making of copies unless permission is obtained from the copyright proprietors. Pertinent discussion including author’s closure, if any, will be published in the July-August 2013 ACI Materials Journal if the discussion is received by April 1, 2013.

Early-Age Creep of Mass Concrete: Effects of Chemical and Mineral Admixturesby Sergio Botassi dos Santos, Luiz Carlos Pinto da Silva Filho, and João Luiz Calmon

binder content.1-9 In other research works, however, priority was given to maintaining the unchanged water-cement ratio (w/c)10-12 and workability, while all other parameters involved in creep could be varied. Few studies have actually focused on the direct effect of admixtures on creep with at least a constant volume of paste.13-15 It is herein intended to main-tain an unchanged material mixture proportion, even with the inclusion of chemical or mineral admixtures, to avoid any change in basic concrete mixture parameters (volume of paste, cement and water content, and aggregate propor-tion) due to mixing with admixtures, which may influence the results. Besides evaluating the real influence of admix-tures on creep, this experimental strategy has the advantage of being able to extrapolate these effects to similar situations because what is actually evaluated is the interference of material added to concrete in creep and not a change in the mixture proportion.

RESEARCH SIGNIFICANCEAttaining better knowledge of mass concrete creep

behavior, using extensive and reliable laboratory data, will enable designers to define mixture proportions considering the potential effects of chemical and mineral admixtures’ contents on creep. This is an important issue to ensure long-term serviceability of massive structures. Such an improve-ment in predicting creep reduces empiricism and permits a more rational use of materials for the manufacture of mass concrete that meets the minimum project requirements.

EXPERIMENTAL INVESTIGATIONFor a more realistic evaluation of the effect of admixtures

on creep, as quoted in the Introduction, the mixture propor-tion (w/c, water content, paste volume, and fixed total aggre-gate/binder ratio) was maintained unaltered throughout the experimental program with variation in only the type of plasticizer and mineral admixture relative to the reference mixture proportion.

To maintain the mixture proportion fixed, part of the reference amount of cement was replaced by the respec-tive mineral admixture studied while still maintaining the amount of water fixed to ensure that the main prompters of creep in the mixture were kept unaltered and, hence, focusing the study on only the effect of mineral admixtures. With regard to the plasticizer admixtures, the concern was to remove part of the mixing water to compensate for the

Some advances in concrete chemical and mineral admixtures are well-known, especially those related to rheology in the fresh state, improvements in the binder-matrix microstructure, and mechanical behavior. When it comes to the effects on creep, however, there are still several issues that are not well understood, especially during the early ages at loading. To help fill this gap, an experimental program was developed considering a laboratory test to monitor creep under controlled environmental conditions, altering the composition of the reference concrete with the inclusion of sepa-rate admixtures of calcined clay, metakaolin and blast-furnace slag, and lignosulfonate- and naphthalene-based chemical admix-tures, keeping binder paste content unchanged to make it possible to evaluate the real effect on basic creep. The obtained results show a significant change in creep during the early ages at loading with the inclusion of the afore-cited materials, especially at the age at loading of 1 day.

Keywords: chemical admixtures; creep; early ages; mass concrete; mineral admixtures.

INTRODUCTIONThe prediction of creep for the analysis of thermomechan-

ical problems in mass concrete structures is of fundamental importance because the delayed effect of strains caused by creep reduces the internal stress history in the structure and therefore influences the appearance of cracks of thermal origin. Typically, these problems are significant in large structures (dams, bridges, and large foundations) in which the structural parts have high concrete content, or even in other concrete structures in which there are high amounts of binder—that is, high-strength concrete.

The use of plasticizing and mineral admixtures in mass concrete can help combat this thermal phenomenon because they contribute to the optimization of the mixture to obtain a higher resistance, lower binder consumption, and an enhanced rheology of the fresh concrete. Normally, however, when it comes to the effect on creep, no specific concern is known to be taken in adjusting the mixture proportion to meet this need. This is due to the relatively little knowledge of the actual and potential effects of mineral admixtures and plasti-cizers on general creep behavior—especially their combined effect—and also on creep during early loading ages.

A survey conducted by the authors of this paper revealed that in many of the research works which, until now, had attempted to evaluate the effects of admixtures on creep, the mixture proportion of concrete was not kept fixed, thus making it difficult to conclude which effect was dominant in the observed change in creep—whether direct from admix-tures or indirect due to changes in the mixture proportion from the use of these materials. Many studies have sought to maintain the compressive strength fixed by adding admix-tures, leading inevitably to a change in the mixture propor-tion of concrete and a change in the volume of paste and


lanic reaction, while calcined clay shows a lower reactivity than the former, but also with pozzolanic effects (common pozzolan with a low calcium oxide content) and, lastly, slag shows a cementing effect (high calcium oxide content). The main characteristics of the cement, plasticizing, and mineral admixtures are summarized in Tables 1 and 2.

The coarse aggregate used had a maximum size of 1.25 in. (32 mm), while the sand was artificial and had the same origin as the coarse aggregate from quartz-biotite schist (quartzite-metamorphic rock). Semiquantitative mineral-ogical analyses showed that the aggregates were composed of 50% quartz, 30% biotite, 10% chlorite, 8% muscovite, and 2% feldspar. Both aggregates had a relative density and absorption capacity of approximately 2.60% and 0.5%, respectively. The proportions were established from a refer-ence mixture proportion of concrete with a possible thermal problem: mass concrete. For the admixtures, the concern was in maintaining the volume of paste fixed for all mixture proportions with a partial replacement of the volume of cement of the reference mixture proportion by admixtures, as shown in Table 3. The percentage of cement replaced by volume adopted for the admixtures was 10%, 30%, and 50% relative to metakaolin, calcined clay, and blast-furnace slag, respectively. On considering the plasticizing admixtures, part of the weight of the mixing water—proportional to the amount of water used in dissolving the plasticizer, consid-ered to be approximately 60% of the total mass of the admix-ture—was deduced. The plasticizing admixture’s content (based on lignosulfonate) and high-range water-reducing admixture (HRWRA) (based on naphthalene) used in rela-tion to the cement mass was 1.0% and 0.5%, respectively.

Mineral and chemical admixture contents in the concrete remained unchanged; otherwise, the experimental investiga-

Sergio Botassi dos Santos is a Researcher and PhD Student in civil engineering at the Federal University of Rio Grande do Sul (UFRGS), Porto Alegre, Rio Grande do Sul, Brazil. He received his MS in civil engineering with an emphasis on structures from the Federal University of Espírito Santo (UFES), Vitória, Espírito Santo, Brazil. His research interests include the concrete technology area, including the numerical modeling of behavior, and he is a specialist in thermal stress analysis in mass concrete.

ACI member Luiz Carlos Pinto da Silva Filho is a Professor and Head of the Graduate Program on Civil Engineering at UFRGS. He received his PhD in civil engineering from the University of Leeds, Leeds, UK. His research interests include various areas of civil engineering materials and structures.

João Luiz Calmon is a Professor at UFES. He received his BSE in civil engineering from UFES; his MSE in industrial engineering from Catholic Pontific University, Rio de Janeiro, Brazil; his PhD in civil engineering from Catalonia Polytechnic Univer-sity, Barcelona, Spain; and his Post-Doctorate degree from Instituto de Ciencias de la Construcción Eduardo Torroja, Madrid, Spain. His research interests include the technology of building materials and experimental and numerical modeling studies of the thermomechanical behavior of mass concrete structures.

inclusion of liquid admixtures, thus maintaining the water content in the mixture and avoiding any modification in the important creep components.

Materials and mixture proportionsPortland cement was used with a minimal amount of lime-

stone filler (cement typical of central Brazil) to obtain a pure binder (without mineral addition) that served as a reference for the study. The main admixtures used were metakaolin, common calcined clay and blast-furnace slag, and lignosul-fonate- and naphthalene-based plasticizing admixtures. The criterion for selecting these items was based on their level of interaction with cement: metakaolin with a high silicon oxide and aluminum content due to their high specific area is known to be more reactive to cement through a pozzo-

Table 1—Physical and chemical compositions of cement and mineral admixtures

Cement Metakaolin Calcined clay Blast-furnace slag

Specific gravity 3.02 2.64 2.60 2.92

Blaine fineness, cm2/g 4900 9840 7930 4850

Loss of ignition, % 5.24 2.06 5.33 0.00

SiO2, % 18.6 50.50 45.00 33.81

Al2O3, % 4.53 38.29 42.75 11.81

CaO, % 57.23 0.55 0.49 42.78

MgO, % 1.53 0.31 0.27 7.13

SO3, % 2.42 * 0.05 1.61

Na2O, % 0.21 0.15 0.65 0.45

K2O, % 0.51 1.49 0.46 0.79

Fe2O3, % 2.45 * * *

*Not measured items.

Table 2—Characteristics of chemical admixtures

Denomination Chemical base Type* Description

Plasticizer(normal water-reducing)

Lignosulfonate A• Addition rate of between 0.25 and 0.50% by weight of cement;

• Specific gravity between 1.21 and 1.25 at 25°C (77°F); and• pH between 8 and 10.

HRWRA Naphthalene F• Addition rate of between 0.60 and 1.50% by weight of cement;

• Specific gravity between 1.18 and 1.22 at 25°C (77°F); and• pH between 7 and 9.

*According to ASTM C494/C494M.


tion would become too long and would lose the comparative focus between the study admixtures and the concrete proper-ties. It is known, however, that mineral and chemical admix-tures’ content interferes in creep, as verified in the creep results obtained by Brooks.12 The hydration mechanism of the binder matrix tends to change depending on the mineral and chemical admixture contents and types, which influence concrete properties.

Properties studied and test proceduresBasic creep was estimated and then deduced from the

effect of autogenous shrinkage obtained from independent tests according to the Brazilian Standard NBR-8224,16 using pairs of cylindrical specimens 5.91 in. (150 mm) in diameter and 11.82 in. (300 mm) in height. For the creep test, the load was applied at 1, 3, and 7 days. Tests were also carried out to determine the compressive strength and modulus of elas-ticity based on the Brazilian Standards—NBR-1282117 and NBR-8522,18 respectively—at ages of 1, 3, 7, and 28 days, carried out on specimens similar to those used in creep and shrinkage determination. The following strength test results are to keep loading at 40% of the compressive strength propor-tionally fixed throughout the execution of the creep test.

To monitor the strain during the creep and autogenous shrinkage tests, strain gauges with optical fiber sensors encapsulated in metal cylinders and embedded in specimens for increased reliability and accuracy of the results were used, as shown in Fig. 1. Strain gauges with optical sensors present a high precision similar to the Carlson19-type elec-trical sensor; in addition, their customization by the manu-facturer for more specific research studies is possible. The tests were monitored for approximately 90 days.

A total of six different mixtures were studied: three with mineral admixtures, one with plasticizer, one with HRWRA, and one without admixtures (reference), totaling 66 tests and 192 test specimens that were cast.

After casting, the specimens were kept in a moisture chamber under a controlled environment (temperature: 23°C ± 2°C [73.4°F ± 3.6°F] and relative humidity: 97.5% ± 2.5%) until the age of onset of each test. The basic creep and autoge-nous shrinkage tests were carried out on sealed samples under a controlled temperature and humidity (23°C ± 2°C [73.4°F ± 3.6°F] and 50% ± 5%, respectively), as shown in Fig. 2.

RESULTS AND DISCUSSIONCompressive strength

The individual results of the compressive strength and modulus of elasticity for each concrete mixture showed a low

dispersion from the average with a coefficient of variation of less than 7% in most of the results. Thus, the authors opted to present the mean values as summarized in Fig. 3 and 4, respectively. In the figures, the dashed-line curve represents results obtained for mixtures with mineral admixtures, while the continuous line represents the measurements obtained for concrete mixtures with chemical admixtures.

From the plot given in Fig. 3, it was verified that all admixtures delayed the gain in compressive strength at the test age of 1 day compared to the reference concrete (without admixtures). This effect is associated with the delay in the formation of hydrated compounds of the binder matrix at early ages caused by the addition of mineral admixtures and plasticizer in concrete mixtures. From the test age of 1 day, however, the rate of increase in the compressive strength of the studied mixture proportions was found to be higher than that observed for the reference concrete, except for concrete with HRWRA. This is reflected in the creep phenomenon

Table 3—Concrete mixture proportions

Rf Mk Cc Bs Sp Pl

Admixture description Reference—without admixture Metakaolin Calcined clay Blast-furnace slag HRWRA Plasticizer

Cement, lb/ft3 (kg/m3) 30.7 (492) 27.7 (443) 21.5 (345) 15.4 (246) 30.7 (492) 30.7 (492)

Admixture, kip/ft3 (kg/m3) 0 2.8 (43) 8.0 (128) 15.0 (240) 3.1 (4.9) 1.5 (2.5)

Water, kip/ft3 (kg/m3) 14.4 (231) 14.4 (231) 14.4 (231) 14.4 (231) 14.2 (229) 14.2 (229)

Artificial sand, kip/ft3 (kg/m3) 29.7 (476) 29.7 (476) 29.7 (476) 29.7 (476) 29.7 (476) 29.7 (476)

Coarse aggregate, kip/ft3 (kg/m3) 67.4 (1079) 67.4 (1079) 67.4 (1079) 67.4 (1079) 67.4 (1079) 67.4 (1079)

Paste volume, % 40.4 40.4 40.4 40.4 40.4 40.4

Air, % 0.9 1.5 1.0 0.6 0.7 1.2

Slump, in. (mm) 3.7 (95) 2.8 (70) 2.2 (55) 3.3 (85) 5.5 (140) 4.3 (110)

Fig. 1—Optical strain gauge prepared for casting and its measuring apparatus.

Fig. 2—Preparation of test specimen (sealing) and load application.


cially at the test ages of 1 to 7 days. Nonetheless, at the age of 28 days, the dispersion was lower. This is mainly due to the effects of chemical and mineral admixtures, which interfere in the rate of hydration and formation of cementi-tious compounds, especially at early ages when the concrete shows more viscoelastic behavior. This effect is reduced at later ages when the elastic behavior of solid materials begins to prevail and also due to concrete mixtures having the same aggregate proportion in relation to the binder matrix, inde-pendent of the admixtures.

The reduction in the modulus of elasticity at very early ages, as observed in this study, may bring unde-sirable consequences in mass concrete structures due to the lower concentration of compressive stress benefi-cial in arresting tensile stresses arising from the thermal problem. However, a closer look at the behavior of the compressive strength of mixtures with admixtures at more advanced ages shows a tendency toward an increase compared to the concrete without admixtures, which helps combat the appearance of thermal cracks.

In general, it was observed that the experimental results obtained from the modulus of elasticity tests performed at 28 days were lower than the expected values estimated directly from the compression test results, in accor-dance with the procedure recommended in the Brazilian NBR-6118 Standard.20 This is possibly due to the high paste content in the study mixture proportions—approxi-mately 40%—as seen in Table 4, which contributed to the reduction of the concrete modulus because the hardened paste usually has lower hardness than the aggregate phase. In the long term, this difference in the elasticity modulus can contribute to reducing tensile stress, which induces thermal cracking in concrete.

Presentation of strain resultsThe test results of autogenous shrinkage and basic creep

are presented as the strain rate over time from the fit of a loga-rithmic curve of the strain values obtained from the tests, as illustrated in Fig. 5 and 6. The values of Fj and Rs, presented in Fig. 5 and 6, are equivalent to the multiplier coefficient of the logarithm of time, as shown in Eq. (1) for autogenous shrinkage and Eq. (2) for the basic creep, respectively.

( ) . ln( )s st R t Bε = + (1)

0 0( ) . ln( )jJ t t F t t B− = − + (2)

at early ages because the hydration rate of cementitious composites with admixtures was significantly different from the reference mixture, besides the effect of admixtures on the microstructure of cementitious matrixes—more pore refine-ment and enhanced bonding between the hydrated particles.

Young’s modulus of elasticityAlthough the mixture proportion—paste volume,

amount of water, and total volume of aggregates—was kept unchanged, the behavior of the modulus of elasticity varied significantly between the studied mixtures, as can be perceived from the results presented in Fig. 4, espe-

Fig. 3—Average behavior of concrete mixtures on compres-sive strength.

Fig. 4—Average behavior of concrete mixtures on Young’s modulus of elasticity.

Table 4—Comparison of autogenous shrinkage ratio of concrete mixtures

Mixture Symbol Autogenous shrinkage ratio, microstrain/ln(t) Increase by reference, %

Reference Rf 55.7 —

With metakaolin Mk 100.0 80

With calcined clay Cc 91.7 65

With blast-furnace slag Bs 114.7 106

With HRWRA Sp 39.9 –28

With plasticizer Pl 37.4 –33


The symbols es(t) and J(t – t0) represent the autogenous shrinkage and basic creep, respectively, at any age t and the start of loading t0. The coefficient B in Eq. (1) and (2) is not considered in the results analyzed because it does not affect the strain rate over time but is used simply to fit the curves.

The logarithmic function was chosen to best represent the increasing behavior of the linear strain rate in the observed experimental creep-time and shrinkage-time results according to the U.S. Bureau of Reclamation21 since the 1960s. All adjusted logarithmic curves achieve a coefficient of determination above 0.81, hence demonstrating good agreement with the test results. The representative units of these rates are associated with the logarithm and are not so simple to interpret. However, it is known that the higher they are, the more significant the effects of strains resulting from autogenous shrinkage and creep. These rates also have the advantage of not only permitting the authors to represent a set of results through a single value but also to estimate the rate of increase of strain of concrete at any age.

This study did not aim to compare prediction models, although the authors have already published a study22 pointing out deviations in the values presented by most models assessed based on test results, mainly when mineral and chemical admixtures are used. This situation shows the need to adjust creep prediction models deriving from the addition of these materials to concrete, as proposed by Brooks12 and commented on by Botassi et al.22

Autogenous shrinkageThe rates of autogenous shrinkage for the studied mixtures

are summarized in Table 4 and Fig. 7. For a clear understanding of the plot in Fig. 7, the columns were hatched in different manners, thus permitting the authors to distinguish the refer-ence mixture (filled hatches) from the concrete mixtures with mineral admixtures (diagonal hatched lines) and those with plasticizing admixtures (hatched horizontal lines).

The autogenous shrinkage of concrete mixtures with mineral admixtures was higher than that of the reference concrete with an increase of over 100% for the blast-furnace slag mixture. This substantial increase is possibly due to the higher fineness of the binder in mixtures with admixtures compared to the reference concrete (without mineral admix-tures), hence promoting a greater internal water consump-tion during the hydration reactions, which favors autogenous shrinkage. This significant increase can also be associated with the reduction in micropore size in the cement matrix, thus resulting in very high surface tensions, which hinders the diffusion of water through matrix voids. The fact that blast-furnace slag shows higher autogenous shrinkage—although it does not have the highest fineness among the studied mineral admixtures—may be associated with the late water consump-tion resulting from capillary voids because its cementitious reaction is slower. However, this tendency requires further and more numerous tests to be statistically confirmed.

Meanwhile, for the concretes with plasticizer and HRWRA, observations showed a reduction of up to 30% in the autog-enous shrinkage. This phenomenon can be explained by the excess water which, in the reference mixture, was mainly used to gain plasticity and help in the hydration of the compounds, while in the presence of the studied chemical admixtures; besides being used for the hydration reactions, this water served to occupy more voids in the cement paste, thus causing an increase in the macro- and mesopores. These pores act as an internal moisture reservoir that reduces

shrinkage and the capillary forces at the expense of larger voids in the paste. Another explanation may be a possible reduction of the surface tension of capillary water caused by the presence of the HRWRA, which entails a reduction on capillary suction and adsorption, facilitating moisture diffusion in the binder matrix while reducing the internal stresses induced by water particle movement. The reduction of autogenous shrinkage in the presence of plasticizer was observed in some studies reported by Collepardi.23

Basic creepThe basic creep rates for the studied mixtures are summa-

rized in Tables 5 and 6 and Fig. 8. Table 6 was assembled from the results in Table 5 to determine the percentage increase (positive value) or decrease (negative value) in the creep of mixtures with admixtures compared to the reference concrete. At the age of loading of 1 day, the creep measured

Fig. 5—Representation of obtained test results of autogenous shrinkage.

Fig. 6—Representation of obtained test results of basic creep.

Fig. 7—Autogenous shrinkage ratio for concrete mixtures.


hydrated) layers, which therefore increases the potential for creep during the early ages.

At a loading age of 3 days, only concrete with HRWRA presented creep values higher than the reference concrete (approximately 20% higher); the remaining mixtures showed a reduction of up to 37% (with blast-furnace slag). This reversal in behavior observed in most samples with admix-tures at the age of 3 days of loading, where there is still a significant effect of viscoelasticity, may be associated with the substantial increase in the stiffness of concrete with admixtures, as can be observed from the higher growth trend in the compressive strength with age of mixtures with admixtures (Fig. 3). Compressive strength estimates obtained for samples with admixtures achieved an increase of 140 to 280% at 1 to 3 days, while the reference concrete did not exceed a 71% increase. The mixture with the lowest increase in strength with age was the one with HRWRA, which possibly explains why it was the only one with creep higher than the reference concrete at a loading age of 3 days.

At a loading age of 7 days, all mixtures with admixtures achieved a creep value lower than that of the reference concrete. The largest reduction was observed in the blast-furnace slag mixture (–77% reduction), while the smallest was the mixture with HRWRA (–6%). One explanation for this general reduction in basic creep may be linked to a higher gain in compressive strength compared to that of the reference concrete for all mixtures with admixtures, which indirectly implies a better formation of the binder matrix and greater stiffness of the concrete.

In general, concrete mixtures with admixtures showed a reduced creep trend relative to the reference mixture, the higher the age at which loading was started (t0). This behavior can cause a significant interference in the stress state of mass concrete structures, as discussed in the work of Botassi et al.24 Mass concrete structures, when newly cast, generate heat, which promotes beneficial compressive stresses at

in concrete with admixtures achieved a substantial increase, reaching a value exceeding four times the creep observed in the reference concrete for concrete with blast-furnace slag addition. As for concrete with chemical admixtures, the mixture with HRWRA showed the highest increase over the reference concrete by more than 140%. This phenomenon may be associated with the low hydration rate of cementi-tious composites of mixtures with admixtures during the first 24 hours, significantly enhancing the viscous behavior of concrete compared to the elastic modulus, as can be seen in the results of the elastic modulus of concrete with admixtures at the age of 1 day (lower than the reference concrete; Fig. 4). Another fact that corroborates this large increase in creep is due to the microstructure of the concrete mixtures with admixtures which, in general, shows a higher binder matrix pore refinement compared to the reference concrete (without admixtures). Thus, there is a higher retention of water in the micropores and between the C-S-H (calcium-silicate-

Table 5—Basic creep ratio of concrete mixtures

Mixture Symbol

Basic creep ratio, microstrain/ksi.ln(t0) (microstrain/MPa.ln(t0))

t0 = 1 day t0 = 3 days t0 = 7 days

Reference Rf 141.4 (20.51) 82.5 (11.96) 72.3 (10.49)

With metakaolin Mk 315.2 (45.71) 80.0 (11.61) 29.3 (4.25)

With calcined clay Cc 223.1 (32.36) 57.7 (8.37) 32.3 (4.68)

With blast-furnace slag Bs 588.8 (85.40) 51.6 (7.49) 16.8 (2.44)

With HRWRA Sp 349.2 (50.65) 99.9 (14.49) 67.8 (9.83)

With plasticizer Pl 236.4 (34.29) 79.5 (11.53) 54.8 (7.95)

Table 6—Comparison of basic creep ratio of concrete mixtures

Mixture Symbol

Increase of basic creep ratio by reference

t0 = 1 day, % t0 = 3 days, % t0 = 7 days, %

Reference Rf — — —

With metakaolin Mk 123 –3 –59

With calcined clay Cc 58 –30 –55

With blast-furnace slag Bs 316 –37 –77

With HRWRA Sp 147 21 –6

With plasticizer Pl 67 –4 –24

Fig. 8—Basic creep ratio of concrete mixtures.


early ages and, as they cool, tension stresses. Assuming only the observed creep behavior in mixtures with admixtures, these stresses induce a higher risk of thermal cracking. This is because the beneficial effect of the compressive stresses tends to be mitigated by higher creep levels in mixtures with admixtures during the early hours of casting, as indicated in the tests, and the undesirable tensile stresses are combated with less intensity due to low creep levels compared to the reference concrete. In addition, there are also high values of autogenous shrinkage for concrete with mineral admix-tures, which enhance the risks of cracks other than those from thermal origin. It is important to emphasize, however, that this afore-reported effect on the thermal problem refers only to the isolated and direct result of admixtures on creep and their possible consequences on mass concrete. Hence, on confirming that the effect of these materials on concrete mixtures tends to reduce the amount of binders, the reduc-tion in the w/c, among other beneficial effects, can balance the aforementioned negative aspects to the extent of making admixtures a great ally in combating thermomechanical problems in mass concrete, as are usually considered.

As far as creep interference in the thermal problem is concerned, it is noteworthy to mention that the creep tests in this study were carried out using test specimens under compression but, on the other hand, thermal cracking occurs in the tensile stress condition. Because there is usually higher creep deformation under tensile stress1 for the same level of load—because of the lower mechanical strength of concrete under tensile stress—the consideration of results derived from creep tests performed under compression provides safer results. The adoption of lower deformability for the simu-lation of the thermal problem leads to higher tensile stress values than the ones that will tend to develop in real cases. However, it is also known that mineral admixtures contribute to a natural increase in tensile strength. So, it is no wonder that tensile creep for mixture proportions with mineral and chemical admixtures is lower than the reference concrete, which interferes in the occurrence of thermal cracking.

FURTHER RESEARCHIt is the intention of the authors to complement the present

experimental program to evaluate the combined effects of the studied admixtures, given that the simultaneous use of these products is a common practice. However, the experi-mental program presented in this paper was of paramount importance, owing to the fact that the separate behavior of each mineral and plasticizer admixture could be understood separately and also verify whether the superposition effects could be considered valid when analyzing the combined effect of admixtures. A numerical simulation-based evalu-ation on how these effects may in fact intervene in the ther-momechanical problem in mass concrete structures still remains, however, considering not only the effect of admix-tures on creep but also on the main parameters involved in the problem (heat generation in concrete, thermal and mechanical properties), thus obtaining the stress levels in the structures and their cracking probabilities.

As a future study subject, it is suggested that the effects of mineral and chemical admixtures be assessed for long-term creep (180 to 365 days) so deformation behavior when using these products can be evaluated in the long run and the effects on structures subject to long-term loads, such as bridges and dams, can be estimated. It is also recommended that the effects of mineral and chemical admixtures on creep

be assessed for different contents and types of cement and concrete, in which creep can directly or indirectly interfere in the tensile and deformational state of structures.

SUMMARY AND CONCLUSIONSThe direct effect of the mineral and chemical admixtures

studied on creep was shown to be highly significant, espe-cially at the age of loading of 1 day. In this experimental program, there was the concern of maintaining unchanged mixture proportions, even with the addition of admixtures. This was the path chosen because if a different experi-mental strategy were to be chosen—for example, keeping the compressive strength constant for different mixtures—it would be necessary to change mixture proportions. Changes in mixture proportions would make it more difficult to monitor the isolated effects of the use of admixtures in creep behavior because these effects would be superimposed by the combined indirect effects of varying important factors that are known to intervene in creep, such as paste content and the total amount of water in the mixtures.

The autogenous shrinkage test was carried out to be deduced from the measured creep test. A large increase in shrinkage was observed for tests on mixtures with mineral admixtures compared to the reference concrete, reaching a peak value of 106% for blast-furnace slag mixtures. Mean-while, for the plasticizing admixtures, an average reduction in autogenous shrinkage of 30% was observed.

In summary, the obtained results of creep and the main conclusions can be summarized as follows:

1. At the loading age of 1 day, mixtures with admix-tures showed a significantly higher creep than the reference concrete (without admixtures), ranging from a 58% increase for mixtures with calcined clay up to a 316% increase for blast-furnace slag. This effect is due to the highly viscous behavior of concrete at very early ages, exacerbated by the delayed hydration of the binders and pore refinement of the matrix when the admixtures are added.

2. At the loading age of 3 days, the mixture with HRWRA achieved a creep 21% higher than the reference concrete. The remaining mixtures showed a reduced creep of up to 37% (for blast-furnace slag). The observed creep reversal is due to the significant improvement in matrix stiffness (1 to 3 days) for mixtures with admixtures, notwithstanding the high pore refinement resulting from the use of these addi-tional materials in cement.

3. At a loading age of 7 days, all mixtures achieved creep values lower than the reference concrete, reaching levels of up to a 77% reduction for the blast-furnace slag mixture. This general reduction in creep is a consequence of the high increase in the stiffness, similar to that commented on at the age of 3 days, and a likely improvement in the aggregate-cement paste interface transition zone.

4. The trend toward increased creep during early ages of testing and its reduction at ages greater than 3 days confirm that admixtures contribute to the enhancement of thermome-chanical problems in mass concrete when only the direct effect on creep is analyzed. However, it is important to highlight that there are other significant effects promoted by admix-tures in mass concrete, which can contribute to reducing the risk of thermal cracking with the possibility of reducing the cement consumption and increasing the mechanical strength.

5. Granulated blast-furnace slag had the highest influence on creep behavior. The concrete mixture containing this mineral admixture showed creep values four times higher


than the reference concrete at a loading age of 1 day and a reduction of 77% in creep at 7 days, as previously mentioned. On the other hand, the HRWRA mixture achieved the highest increase in creep, with a 2.5-fold increase at an age of loading of 1 day compared to the reference concrete but a milder reduction of only 6% from the seventh day of loading.

ACKNOWLEDGMENTSThe authors are grateful to Laboratório de Concreto de Furnas Centrais

Elétricas S.A., Brazil, for their unconditional support throughout the experimental program; to Agência Nacional de Energia Elétrica (ANEEL) for the financial support to the research project; and to G. Sensors, repre-sentative of Fiber Sensing Products in Brazil for her partnership with the researchers to customize the optical extensometers—an application precursor in creep tests. The authors are also grateful to the civil engineers M. Alexandre, A. Neiry, A. Liduário, E. Gambale, F. de Lima, F. Mamede, and A. de Castro, as well as the technicians L. Matiazzo, J. Bonifácio, G. Ramos, and Á. Donizete for their technical and operational support of the project.

NOTATIONBs = concrete mixture with blast-furnace slagCc = concrete mixture with calcined clayC-S-H = calcium-silicate-hydratedes(t) = autogenous shrinkage strainFj = basic creep ratioHRWRA = concrete mixture containing HRWRAJ(t – t0) = creep compliance functionMk = concrete mixture with metakaolinPl = concrete mixture with plasticizerRf = reference concrete (without admixtures)Rs = autogenous shrinkage ratiot = age of concretet0 = age of concrete from start of loading during creep test

REFERENCES1. Neville, A. M., Creep of Concrete: Plain, Reinforced and Prestressed,

first edition, North-Holland Publishing Company, Amsterdam, the Nether-lands, 1970, 622 pp.

2. Brooks, J. J., and Neville, A. M., “Creep and Shrinkage of Concrete as Affected by Admixture and Cement Replacement Materials,” Creep and Shrinkage of Concrete: Effect of Materials and Environments, SP-135, M. A. Daye and C. C. Fu, eds., American Concrete Institute, Farmington Hills, MI, 1992, pp. 19-36.

3. Rixom, R., and Mailvaganam, N., Chemical Admixtures for Concrete, third edition, E&FN Spon, London, UK, 1999, 456 pp.

4. Alexander, K. M.; Bruere, G. M.; and Ivanusec, I., “The Creep and Related Properties of Very High-Strength Superplasticized Concrete,” Cement and Concrete Research, V. 10, No. 2, Mar. 1980, pp. 131-137.

5. Brooks, J. J., and Jiang, X., “The Influence of Chemical Admixtures on Restrained Drying Shrinkage of Concrete,” Superplasticizers and Other Chemical Admixtures in Concrete, SP-173, V. M. Malhotra, ed., American Concrete Institute, Farmington Hills, MI, 1997, pp. 249-265.

6. Luther, M. D., and Hansen, W., “Comparison of Creep and Shrinkage of High-Strength Silica Fume Concretes with Fly Ash Concretes of Similar Strengths,” Fly Ash, Silica Fume, Slag and Natural Pozzolans in Concretes—Proceedings of the 3rd International Conference, SP-114,

V. M. Malhotra, ed., American Concrete Institute, Farmington Hills, MI, 1989, pp. 573-591.

7. Timusk, J., and Ghosh, R. S., “Creep of Fly Ash Concrete,” ACI JOURNAL, Proceedings V. 78, No. 5, Sept.-Oct. 1981, pp. 351-357.

8. MacGregor, I. D., “A Comparison of Mechanical Properties of Hong Kong Medium and High-Strength NPC and Fly Ash Concretes,” CANMET/ACI Sixth International Conference on Fly Ash, Silica Fume, Slag and Natural Pozzolans in Concrete, Bangkok, Thailand, May 31-June 5, 1998, pp. 129-158.

9. Nasser, K. W., and Al-Manaseer, A. A., “Creep of Concrete Containing Fly Ash and Superplasticizer at Different Stress/Strength Ratios,” ACI JOURNAL, Proceedings V. 83, No. 4, July-Aug. 1986, pp. 668-673.

10. Brooks, J. J., “How Admixtures Affect Shrinkage and Creep,” Concrete International, V. 21, No. 4, Apr. 1999, pp. 35-38.

11. Ramachandran, V. S., Concrete Admixtures Handbook: Properties, Science and Technology, second edition, Noyes Publications, Park Ridge, NJ, 1995, 626 pp.

12. Brooks, J. J., “Elasticity, Creep and Shrinkage of Concretes Containing Admixtures,” The Adam Neville Symposium: Creep and Shrinkage—Struc-tural Design Effects, SP-194, A. Al-Manaseer, ed., American Concrete Institute, Farmington Hills, MI, 2000, pp. 283-360.

13. Buil, M., and Acker, P., “Creep of a Silica Fume Concrete,” Cement and Concrete Research, V. 15, 1985, pp. 463-466.

14. Li, H.; Wee, T. H.; and Wong, S. F., “Early-Age Creep and Shrinkage of Blended Cement Concrete,” ACI Materials Journal, V. 99, No. 1, Jan.-Feb. 2002, pp. 3-10.

15. Akkaya, Y.; Ouyang, C.; and Shah, S. P., “Effect of Supplementary Cementitious Materials on Shrinkage and Crack Development in Concrete,” Cement and Concrete Composite, V. 29, No. 2, 2006, pp. 117-123.

16. Brazilian Technical Standards Association, “Concrete—Determination of Creep: NBR-8224,” Rio de Janeiro, Brazil, 1983, pp. 1-10. (in Portuguese)

17. Brazilian Technical Standards Association, “Concrete—Preparation in Laboratory—Procedure: NBR-12821,” Rio de Janeiro, Brazil, 2009, pp. 1-5. (in Portuguese)

18. Brazilian Technical Standards Association, “Concrete—Determi-nation of the Elasticity Modulus by Compression: NBR-8522,” Rio de Janeiro, Brazil, 2008, pp. 1-16. (in Portuguese)

19. Botassi, S. S.; Calmon, J. L.; Silva Filho, L. C. P.; Liduário, A. S.; and Gambale, E. A., “Comparative Evaluation Behavior of Optical and Resis-tance Electric Strain Gages Embedded in Samples of Concrete,” Proceed-ings of the 53rd Brazilian Concrete Congress, IBRACON, Fortaleza, Brazil, 2010, pp. 1-15. (in Portuguese)

20. Brazilian Technical Standards Association, “Design of Structural Concrete—Procedure: NBR-6118,” Rio de Janeiro, Brazil, 2007, pp. 1-221. (in Portuguese)

21. U.S. Bureau of Reclamation, “Creep of Concrete under High Inten-sity Loading,” Concrete Laboratory, Report No. C-820, Denver, CO, 1956, 395 pp.

22. Botassi, S. S.; Calmon, J. L.; Silva Filho, L. C. P.; and Gambale, E. A., “The Effects of Mineral and Chemical Admixtures on Creep of Concrete Using Prediction Models,” Proceedings of the 52nd Brazilian Concrete Congress, IBRACON, Florianópolis, Brazil, 2011, pp. 1-13. (in Portuguese)

23. Collepardi, M. M., “Water Reducers/Retarders,” Concrete Admix-tures Handbook—Properties, Science and Technology, second edition, V. S. Ramachandran, ed., Noyes Publications, Park Ridge, NJ, 1995, pp. 286-409.

24. Botassi, S. S.; Calmon, J. L.; Silva Filho, L. C. P.; and Andrade, M. A. S., “Basic Creep Prediction on Database of the Furnas Concrete Laboratory: Preliminary Studies,” Proceedings of the 49th Brazilian Concrete Congress, IBRACON, Bento Gonçalves, Brazil, 2007, pp. 1-12. (in Portuguese)


Title no. 109-M53


ACI Materials Journal, V. 109, No. 5, September-October 2012.MS No. M-2011-224.R1 received November 30, 2011, and reviewed under Institute


Proposed Flexural Test Method and Associated Inverse Analysis for Ultra-High-Performance Fiber-Reinforced Concreteby Florent Baby, Benjamin Graybeal, Pierre Marchand, and François Toutlemonde

stress-strain constitutive law, an inverse analysis is necessary to determine the uniaxial tensile behavior.

Analytical inverse analyses for four-point flexural tests on UHPFRC or high-performance fiber-reinforced cementi-tious composites (HPFRCCs) have been developed by many researchers (AFGC-SETRA 2002; Ostergaard et al. 2005; Kanakubo 2006; Qian and Li 2008; Rigaud et al. 2011) with some success. The AFGC-SETRA (2002) simplified inverse method for thin UHPFRC elements, which engages the applied load and the midspan deflection, is based on the equilibrium of moments and forces in a sectional analysis. This method assumes a bilinear curve for the UHPFRC constitutive law and knowledge of the initiation point for nonlinear behavior. However, the nonlinear behavior initia-tion point value obtained through four-point bending tests cannot be directly used due to a significant size effect (Chanvillard and Rigaud 2003; Frettlöhr and Reineck 2010), thus inducing difficulties for the analysis of experimental results concerning the loss of linearity.

Analytical methods developed by Kanakubo (2006) and Qian and Li (2008), which use a sectional analysis similar to that developed by Maalej and Li (1994), are based on a simplified tensile elastic-plastic curve. The Qian and Li (2008) procedure, which has been developed for the quality control of strain-hardening cementitious composite (SHCC), requires only the knowledge of the maximum applied load and the corresponding load-point deflection. The Kanakubo (2006) method is based on the measurement of the applied load and the curvature in the constant bending moment zone. These simplified inverse methods are not able to predict the real tensile stress-strain curve when considered in terms of a point-by-point curve. Moreover, the assumption of uniform stress distribution along the tensile height can induce a strength overestimation.

The Rigaud et al. (2011) inverse analysis method uses the experimentally captured bending-moment-versus-midspan-deflection response, which is converted into a bending-moment-versus-curvature response through an equation that relates the midspan deflection of the prism to the curvature along the middle third of the span. This equation, which is similar to the one used by Qian and Li (2008), is based on structural elastic mechanics and considered as reasonably

The tensile stress-strain response of ultra-high-performance fiber-reinforced concrete (UHPFRC) is a fundamental constitu-tive property, and reliable knowledge of this response is neces-sary for appropriate application of the tensile-carrying capacity of such advanced cement-based materials. Flexural test methods whose implementation is well-established present a test procedure capable of assessing this property. Nevertheless, these methods provide indirect information and need to be complemented by inverse analysis to quantify the intrinsic tensile behavior of tested materials. Moreover, bias or scatter can be induced when simpli-fied constitutive laws are assumed for the analysis. Flexural tests were completed on multiple types of commercially available UHPFRC. Relying on direct strain measurements, a new inverse analysis method is presented and qualified, compared with an existing simplified method, and also compared with results from direct tensile tests (DTTs). The advantages and limitations of the experimental and analysis methods were derived.

Keywords: bending test; flexural test; inverse analysis; strain distribution; tensile stress-strain response; ultra-high-performance concrete (UHPC); ultra-high-performance fiber-reinforced concrete (UHPFRC).

INTRODUCTIONUltra-high-performance fiber-reinforced concrete (UHPFRC)

is a class of cementitious composite materials designed to exhibit outstanding mechanical and durability properties, including sustained postcracking tensile strength (Richard and Cheyrezy 1995; Naaman and Reinhardt 1996; Behloul 1996; AFGC-SETRA 2002; Chanvillard and Rigaud 2003; Walraven 2009; Toutlemonde and Resplendino 2010; Graybeal 2011). Laboratory tests of structural elements have clearly indicated that UHPFRC components can exhibit tensile mechanical properties far in excess of those expected from conventional concretes and fiber-reinforced concretes (FRCs) (Graybeal 2006a, 2009; Sato et al. 2008; Baby et al. 2010; Bertram and Hegger 2010; Toutlemonde et al. 2010). Specific quantification of these tensile mechanical properties has proven difficult, however, leading to hesitancy among designers considering the engagement of these properties in UHPFRC components within the civil infrastructure. Many researchers have attempted to develop test methods for the assessment of the tensile performance of FRCs. Test methods have included both direct and indirect assessments, including some that have been standardized (RILEM TC162-TDF 2001, 2002). Most of them are based on the definition of a stress-crack-opening law, consistent with design methods of reinforced concrete (RC) and conventional FRC structures. Due to the multiple-fine-cracking behavior of UHPFRC elements, however, a stress-strain approach is more appro-priate. When using four-point flexural tests for identifying a


finite element analysis, but no direct comparison with direct tensile test results has been published.

RESEARCH SIGNIFICANCEThe focus of this research was to optimize an analysis

method for deriving the tensile stress-strain response of UHPFRC from four-point flexural tests. The midspan strain measurement on the specimen tensile face, as captured by two staggered extensometers, is used to obtain the experimental bending-moment-versus-midspan strain on the tensile face response and also to determine the crack localization. Then, a point-by-point inverse analysis is used to derive the tensile stress-strain relationship. Thus, the UHPFRC tensile stress-strain relationship is derived through a method that reduces the reliance on assumed behaviors, thereby increasing the fundamental soundness of the analytically produced results.

PROPOSED METHODConcerning the tensile stress-strain response of UHPFRC,

the easiest way to determine the strain value without making any assumptions is to use a direct measurement. In this test program, two linear variable differential transformers (LVDTs) used as extensometers are applied to the tensile face of each specimen to measure the midspan strain on the tensile face and determine the crack localization. Then, the tensile stress-strain relationship of the tested material is derived from the experimental bending-moment-versus-midspan strain on the tensile face response without assuming the profile of the tensile stress-strain curve.

Determination of crack localizationThe use of a pair of staggered LVDTs allows for simpli-

fied identification of crack localization (as shown in Fig. 1). This setup helps to distinguish the onset of bifurcation of the cracking process with crack localization over one of the gauge lengths while cracking remains diffuse over the other gauge length, as shown in Fig. 2(a).

In some cases, two localized cracks can occur before one main failure crack develops, or the localized crack can be detected by both LVDTs, as shown in Fig. 2(b). For these cases, the crack localization is assumed to correspond to the maximum bending stress.

In some instances, such as the one shown in Fig. 2(c), three steps can be observed:• First step: Displacements measured by both LVDTs

increase;• Second step: One displacement stops increasing; and• Third step: An unloading branch occurs with a

decreasing value for one displacement.In this case, the experimental bending-moment-versus-

midspan strain on the tensile face response can exhibit a long plateau with little increase of the load before reaching the maximum load. During this step, the displacement rate reported by one displacement stops increasing, indicating that the damage is not completely localized. This step could be explained by a very low stress decrease in the “localized crack” combined with the bending configuration, which allows stabilization or a small increase in the load.

Clearly, in some cases, crack localization can occur before reaching the maximum bending stress. This is not addressed within the methods described by Kanakubo (2006), Qian and Li (2008), and Rigaud et al. (2011), each of which assume that crack localization corresponds with the maximum bending stress.

Florent Baby is a Researcher in the Structures Department of the French Institute of Science and Technology for Transport, Development and Networks (IFSTTAR) (formerly the French Central Laboratory of Roads and Bridges [LCPC]), Paris, France. He graduated from Ecole Nationale des Travaux Publics de l’Etat and Master MEGA at Lyon, Lyon, France. His research interests include the behavior of structures made of ultra-high-performance fiber-reinforced concrete (UHPFRC) and advanced cementitious materials characterization.

ACI member Benjamin Graybeal leads the Structural Concrete Research Program for the Federal Highway Administration at the Turner-Fairbank Highway Research Center, McLean, VA. He received his BS and MS from Lehigh University, Bethlehem, PA, and his PhD from the University of Maryland, College Park, MD. He is a member of ACI Committee 239, Ultra-High Performance Concrete. His research interests include structural application of advanced cementitious materials, concrete material characterization, experimental evaluation of highway bridge structures, and nondestructive evaluation techniques.

Pierre Marchand is the Head of the Structural Engineering Unit of IFSTTAR (formerly LCPC). He graduated from École Polytechnique, Palaiseau, France, and École Nationale des Ponts et Chaussées, Champs-sur-Marne, France. His research interests include steel and concrete structures, including UHPFRC.

ACI member François Toutlemonde is the Deputy Head of the Bridges and Structures Department of IFSTTAR (formerly LCPC). He graduated from École Polytechnique and École Nationale des Ponts et Chaussées and is the President of the ACI Paris Chapter. His research interests include high-rate dynamics of concrete structures; structural effects of alkali-aggregate reaction (AAR) and delayed ettringite formation (DEF); and structural applications of high-performance concrete (HPC), fiber-reinforced concrete (FRC), and UHPFRC.

valid for nonlinear behavior. The method is based on the equilibrium of moments and forces in a sectional analysis for each value of curvature, and the corresponding bending moment does not need to assume the profile of the tensile stress-strain relationship. A genuine constitutive point-by-point tensile stress-strain curve is thus derived. Nevertheless, the equation used to convert the deflection into curvature induces an underestimation of the curvature for a given value of the deflection. As a consequence, methods based on this mechanical assumption underestimate the real strain during the pseudo-strain-hardening phase and overestimate the postcracking stress.

A hinge model that takes into account the tensile strain hardening and the influence of the crack localization including the softening (that is, stress versus crack opening) behavior, has been used by Ostergaard et al. (2005) in an inverse analysis procedure to derive tensile mechanical prop-erties from a beam flexural response. The localized defor-mation has to be determined from the actual localization mechanism, specifically from experimental observations for each tested specimen. Tensile properties derived from this procedure agree well with results developed through

Fig. 1—Midspan strain measurement: staggered extenso-meters on tensile face.


Point-by-point inverse analysisThe experimentally captured bending-moment-versus-

midspan strain on the tensile face response is converted into a tensile stress-strain curve through an inverse method appli-cable from elastic loading through crack localization. The stress-strain curve is based on the equilibrium of moments and forces in a sectional analysis for each value of midspan strain on the tension face and the corresponding bending moment. Assumption of the profile of the tensile stress-strain relationship is not required. The main difference with the point-by-point inverse analysis of Rigaud et al. (2011) is the fact that the experimental midspan strain at the extreme tension fiber is directly measured, not derived from a global measurement and a mechanical assumption.

The strain distribution is considered as linear. This assump-tion is acceptable if the UHPFRC has a pseudo-strain-hard-ening behavior in tension. The compressive behavior of UHPFRC is assumed to be linear elastic, which is realistic for this kind of material (Behloul 1996).

For each strain measurement, the position of the neutral axis is determined via the inverse analysis, as detailed from Eq. (1) to (15). The width and height of the specimen are noted b and h, respectively. The variable E is the elastic modulus (obtained either from compressive tests on cylin-ders or derived from the elastic part of the experimentally captured bending-moment-versus-midspan strain on the tensile face response). The height of the zone under tension is anh and F is the curvature. Compressive stresses and strains are considered as negative and tensile stresses and strains are considered as positive.• In the zone under compression

( )2

2

( ) ( )

12

n n

h h

c c nh h

n

N b z dz b E z h dz

hb E

a a= ⋅s ⋅ = − ⋅ ⋅f ⋅ − a ⋅∫ ∫

= − ⋅ ⋅f ⋅ a − ⋅(1)

( )3

3( ) 2 36n

h

c c n nh

hM b z z dz b Ea

= ⋅s ⋅ ⋅ = − ⋅ ⋅f ⋅ + a − a ⋅∫ (2)

• In the zone under tension

( )( )t nz h ze = f⋅ a − (3)

At the extreme tension fiber, et is equal to etf and at the neutral axis, et is equal to zero. Thus

( )0 0

( )tfnh

t t t t t tbN b dz d

ea

= ⋅s e ⋅ = ⋅ s e ⋅ e∫ ∫f(4)

20 0

( ) ( )tfnh

t t t n t t t t tbM b z dz h N d

ea

= ⋅s e ⋅ ⋅ = a ⋅ − ⋅ s e ⋅ e ⋅ e∫ ∫f

(5)

This inverse analysis uses a discretization of the tensile stress-strain relationship (et,i, st,i). The previous equations can be used to consider two successive loading steps in the section: the loading step i and the loading step i + 1. Between these two loading steps, the strain at the extreme tension fiber increases from etf,i to etf,i + 1 and the corresponding stress changes from st,i to st,i + 1. For these two loading steps, there are two different curvatures and two neutral axis positions.

Therefore, at loading step i

,

,0

( )tf i

t i t t ti

bN de

= ⋅ s e ⋅ e∫f(6)

,

, , , 20

( )tf i

t i n i t i t t t ti

bM h N de

= a ⋅ − ⋅ s e ⋅ e ⋅ e∫f

(7)

Fig. 2—Proposed method to detect crack localization with identification of elastic unloading.


has to be deduced. Thus, et,fi + 1 is computed to take into account the off-plane distance (OPD) of the LVDTs

, 1, 1 , 1

, 1 OPDn i

tf i tf i measuredn i

hh+

+ + −+

a ⋅e = × e

a ⋅ +(14)

Finally, starting from etf,i + 1 – measured, etf,i + 1 and then st,i + 1 are determined.

Because the description of the test results is discrete and the inverse method uses a sort of derivative of the moment curve, oscillations of the tensile stress-strain relationship often occur. It has been shown that it can be stabilized by correcting iteration i after calculating iteration i + 1. In prac-tice, it is sufficient to reposition the stress of iteration i by determining a moving average of the following type

( ), , , 1123t i t i t i+s = ⋅s + s × (15)

If the stress does not vary suddenly (which is the case in practice), this correction does not affect the response and leads to much more realistic results. This stabilization operation must be carried out at the end of each iteration to be taken into account in the calculation of the following iterations. From the raw result, a smooth constitutive curve further usable for design can be obtained by using a polyno-mial interpolation or a moving average.

The validation of the proposed model was first established by using a simple self-consistency case, which consists of generating a bending-moment-versus-strain curve by a direct calculation, then verifying that the result obtained with the inverse analysis is similar to the tensile stress-strain relation-ship used in the direct calculation. Moreover, the procedure has been used for the interpretation of a dedicated experi-mental program described in the following.

EXPERIMENTAL PROGRAM: FLEXURAL TESTSSpecimens and parameters

The experimental program included the completion of four-point flexural tests on four sets of UHPFRC specimens and other associated tests, such as direct tension tests, as well as compressive tests aimed at determining the UHPFRC constitutive law in compression. Table 1 provides details on the four sets of specimens, including which tests were completed on each set. The first character of the specimen name indicates the type of UHPFRC material used and the second character indicates the type of curing regime applied.

At loading step i + 1

, 1

, 1

,

, 101 1

,1

( )

( )

tf i

tf i

tf i

it i t t t

i i

t i t t ti

bN d

bN d

+

+

e

++ +

e

e+

f= ⋅ s e ⋅ e =∫f f

⋅ + ⋅ s e ⋅ e∫f

(8)

( ), 1

,

2

, 1 , 1 , 1 , , ,21

21

( )tf i

tf i

it i n i t i t i n i t i

i

t t t ti

M h N M h N

b d+

+ + ++

e

e+

f= a ⋅ + ⋅ − a ⋅

f

− ⋅ s e ⋅ e ⋅ e∫f

(9)

For both the previous equations, the last term can be expressed in discrete form using the trapezoidal method for integral computation, so that Nt,i + 1 and Mt,i + 1 read

( ), 1 ,, 1 , , 1 ,

1 1

12

t i t iit i t i tf i tf i

i i

N N b ++ +

+ +

s + sf= ⋅ + ⋅ ⋅ ⋅ e − e

f f(10)

( )

( )

2

, 1 , 1 , 1 , , ,21

, 1 , 1 , ,, 1 ,2

1

2

it i n i t i t i n i t i

i

t i tf i t i tf itf i tf i

i

M h N M h N

b

+ + ++

+ ++

+

f= a ⋅ + ⋅ − a ⋅

fs ⋅ e + s ⋅ e

− ⋅ ⋅ e − e⋅f

(11)

All parameters at loading step i are considered as already determined. Thus, solving this inverse problem consists of determining the parameters an,i + 1 and st,i + 1 to satisfy mechanical equilibrium in the section

, 1 , 1 0t i c iN N+ ++ = (12)

, 1 , 1 1c i t i i experimentalM M M+ + + −+ = (13)

An option to implement this inverse analysis could be to iterate on an,i + 1 with respect to Eq. (13).

Concerning the tensile strain at midspan, the effect of the possible additional lever arm due to sensor fixation devices

Table 1—Sets of test specimens and UHPFRC material properties

Specimen set UHPFRC

Steel-fiber volumetric percentage

Curing regime

Four-point flexure—

short

Four-point flexure—

long

Direct tensile test

(DTT)—shortDTT—

long

Density, kg/m3 (lb/ft3)

Compressive strength, MPa

(ksi)

Modulus of elasticity, GPa (ksi)

F1A F 2 Steam — X X X2570

(160.4)220

(32.0)61.0

(8840)

F2A F 2 Lab — X X X2545

(158.9)192

(27.9)62.8

(9110)

F1C F 2.5 Steam X X X X2569

(160.4)212

(30.7)60.3

(8740)

B2A B 2.5 Lab X X X —2690

(168.0)213

(30.9)63.9

(9270)


A “1” indicates that the specimen set was subjected to steam treatment after setting at 90°C (194°F) and 95% humidity for 48 hours, while a “2” indicates that the specimen set was held in a standard laboratory environment until the time of testing at 90 days. The intent of the steam treatment is to increase the mechanical characteristics of the concrete and to accelerate attainment of the final maturity of the heat-treated component. For UHPFRC “F,” the steam treatment is part of the usual manufacturing process. All specimens in a particular set were cast from an individual batch of UHPFRC. Three UHPFRC mixtures were engaged in this study (Table 2). The UHPFRC “F” mixtures are effectively the same, aside from the two different volumetric percent-ages of fiber reinforcement. This particular UHPFRC is commercially available in North America. UHPFRC “B” was included in this study so as to engage a different class of UHPFRC. This UHPFRC is commercially available in Europe. The density, compressive strength, and compressive modulus of elasticity for each set of specimens are provided in Table 1.

Two different specimen lengths with corresponding changes in four-point flexural test configuration were tested within the program. “Long” refers to a 431.8 mm (17 in.) long prism with a cross section of 50.8 x 50.8 mm (2 x 2 in.), a span of 355.6 mm (14 in.), and a distance between the upper rollers equal to 101.6 mm (4 in.). “Short” refers to

a 304.8 mm (12 in.) long prism with a cross section of 50.8 x 50.8 mm (2 x 2 in.), a span of 228.6 mm (9 in.), and a distance between the upper rollers equal to 76.2 mm (3 in.). In all cases, the specimens were single-point cast-in pris-matic molds, allowing the UHPFRC to flow along the length of the form.

Loading setup and instrumentationAll bending tests were completed in a four-point flexural

loading configuration. During the test, the load, deflection of the prisms, and midspan strain at the bottom flange (two staggered values) were measured.

The loading control of the test was accomplished by completing the test in a servo-hydraulic load frame. The control signal for all tests was the stroke with the imposed rate equal to 0.25 mm/minute (0.001 in./minute), as recom-mended in AFGC-SETRA (2002).

The two upper load points and the two lower support points were realized using steel rollers that impart no axial restraint on the prism. The blocks above the upper rollers were supported by 51 mm (2 in.) deep solid steel beams that were connected to a spherical bearing, which ensures that all rollers are bearing on the prism during the test. This assembly has to be set on the specimen prior to the start of the test. As a consequence, the influence of the upper block weight (26 kg [57 lb]) is taken into account by an analytical post treatment.

Concerning the measurement of the midspan deflection, a yoke similar to the one recommended by ASTM C1018-97 is used to measure net values exclusive of any extraneous effects due to seating or twisting of the specimen on its supports or deformation of the support system.

Point-by-point tensile stress-strain relationships developed from proposed inverse analysis

In Fig. 3, the average tensile stress-strain relationships obtained from the proposed point-by-point inverse analysis method are presented for each specimen group (with five or six specimens per batch). These curves are obtained by applying the inverse method to the average point-by-point bending-moment-versus-midspan strain on the tensile face curve and then (from the raw results) using a third-degree polynomial interpolation with a strain interval equal

Table 2—UHPFRC mixtures

MaterialUHPFRC

“F-2%,” kg/m3UHPFRC

“F-2.5%,” kg/m3UHPFRC “B,”

kg/m3

Premix 2195 2161 2296

High-range water-reducing admixture

30 29 50

Steel fibersFf = 0.2 mm; Lf = 13 mm

156 195 0

Steel fibersFf = 0.3 mm; Lf = 20 mm

0 0 195

Water 130 128 190

Notes: 1 kg/m3 = 1.685 lb/yd3; 1 mm = 0.039 in.

Fig. 3—Average tensile stress-strain relationships obtained from proposed point-by-point inverse analysis method.


a consequence, a statistical size effect that induces a lower mean value would be expected. Nevertheless, the experi-mental results show the contrary phenomenon (the results for long specimens are higher than for short specimens), which could be explained by the fact that the longer the prism, the more preferential the orientation of the fibers. The effects of fiber size, fiber orientation, and the specimen casting method influence the test results (Spasojevic 2008).

The comparison of results between Specimens F1A-L and F2A-L shows the well-known effect of the steam treat-ment on the mechanical properties (Behloul 1996; Graybeal 2006b).

COMPARISON OF PROPOSED POINT-BY-POINT PROCEDURE WITH SIMPLIFIED METHOD

The point-by-point tensile stress-strain curves derived from the proposed procedure may be useful to appreciate the real tensile postcracking behavior of UHPFRC. Nevertheless, simplified bilinear curves are convenient when considered in terms of design issues or finite element model (FEM) analyses. Thus, average and characteristic bilinear curves can be obtained from the proposed procedure by using a linear interpolation of the postcracking part of the curves resulting from the inverse analysis previously described herein (refer to Fig. 4). The construction of the characteristic bending-moment-versus-midspan strain on the tensile face curve begins by determining, for each strain interval, the mean value of the bending moment (with five or six speci-mens per batch) and the standard deviation. The character-istic point-by-point bending-moment-versus-midspan strain on the tensile face curve is obtained by subtracting the corre-sponding standard deviation from the mean value multiplied by the Student coefficient (Student’s law with 5% quantile) equal to 2.015 for six specimens and 2.132 for five speci-mens. For the characteristic tensile stress-strain curves, the final strain emin-ppt is equal to the minimum of the following strains (emin-ppt-1, emin-ppt-2):• For each specimen group, emin-ppt-1 is the minimum of all

eSpecimen-k-min-ppt, where eSpecimen-k-min-ppt is the strain corre-sponding with an identification of the elastic unloading (if identified) or the strain at the maximum equivalent bending stress for the specimen k.

• For each specimen group, emin-ppt-2 is the strain corre-sponding to an irreversible decreasing of the stress in the stress-strain curve obtained from the inverse analysis of the characteristic bending-moment-versus-strain relationship.

These bilinear curves obtained from the results of the proposed point-by-point method can be compared with those

to 300 mm/m. The final strain eend-ppt of these curves is the minimum of the following strains (eend-ppt-1, eend-ppt-2):• For each specimen group

5 or 6

- -1 - - -1

15 or 6end ppt Specimen k end ppt

k =e = ⋅ e∑ (16)

where eSpecimen-k-end-ppt is the strain corresponding to an identification of the elastic unloading (if identified) or the strain at the maximum equivalent bending stress for the specimen k.

• For each specimen group, eend-ppt-2 is the strain corre-sponding to an irreversible decreasing of the stress in the average stress-strain curve. Before reaching the maximum equivalent bending stress, the load can increase due to the increase of the internal lever arm while the stress at the bottom flange has already begun to decrease. The crack localization has thus occurred before reaching the maximum apparent bending stress.

It should be noted that the procedure that consists of applying the inverse analysis method to the average point-by-point bending-moment-versus-midspan strain on the tensile face curve gives similar results to the procedure that consists of determining the average curve from all tensile stress-strain relationships obtained for each specimen (in taking into account the same final strain eend).

Before reaching crack localization, a difference between the slopes of both curves “Strain 1 versus average strain” and “Strain 2 versus average strain” was observed for many specimens (refer to Fig. 2). This means that the damage is not perfectly homogeneous in the constant bending length and is more important in a particular zone. This phenomenon could induce a dependence of the measured crack localiza-tion strain with the extensometers. For this reason, it can be interesting to compare the average (ea,k) and the minimum (emin,k) of both staggered LVDT measurements at crack localization (refer to Table 3). The deviation between the average strain and the average minimum strain is approxi-mately 20% for the majority of specimen groups, except for Specimen F2A-L, which is 10%. This deviation seems to be quite high. It could be explained by the relatively low number of cracks in the gauge length. Indeed, the maximum gauge length is equal to 5 × Lf-max and the average space between cracks is approximately 0.75 × Lf (Jungwirth 2006). Thus, testing with a longer constant bending moment zone may reduce this deviation.

For long prisms, the length of the constant bending moment zone is more important than for short prisms. As

Table 3—Average strain and average minimum strain corresponding to elastic unloading

Name of specimen group

Average strain (ea,k) (from all prisms k for each specimen group) corresponding to elastic unloading

Average minimum strain (emin,k) (from all prisms k for each specimen group) corresponding to elastic unloading

Average strainStandard deviation

(number of specimens) Average minimum strainStandard deviation

(number of specimens)

B2A-S 0.0097 0.0021 (6) 0.0076 0.0032 (6)

B2A-L 0.0084 0.0023 (6) 0.0069 0.0030 (6)

F1A-L 0.0080 0.0016 (5) 0.0065 0.0032 (5)

F2A-L 0.0054 0.0025 (5) 0.0049 0.0021 (5)

F1C-S 0.0065 0.0029 (6) 0.0048 0.0023 (6)

F1C-L 0.0076 0.0020 (5) 0.0062 0.0027 (5)


derived from a simplified inverse analysis, which directly assumes a bilinear curve for the UHPFRC constitutive law. Concerning the AFGC-SETRA (2002) analysis, the value of the loss of linearity is necessary. Nevertheless, as explained previously, for some specimen groups, the value of the stress at the loss of linearity obtained through four-point bending tests cannot be directly used due to a scale effect. Moreover, the specimen size induces difficulties to measure the curva-ture from two LVDTs fixed with jigs, as indicated in the Kanakubo (2006) method.

For these reasons, the Qian and Li (2008) inverse analysis was chosen as a simplified inverse analysis for the sake of comparison. The global procedure used in the Qian and Li (2008) method to obtain a simplified tensile stress-strain relationship is described in Fig. 5. By conducting parametric

studies based on a flexural behavior model, “master curves” can be constructed in terms of:• Tensile strain capacity with respect to deflection

capacity; and• Normalized modulus of rupture (MOR) (MOR =

6Mmax/bh2) or maximum equivalent flexural stress by effective tensile strength ste with respect to tensile strain capacity.

For the tensile-strain-capacity-versus-deflection-capacity master curve, as compared with the Qian and Li (2008) method, the calculation of deflection corresponds to the midspan deflection (that is, not the load-point deflection) and the stress distribution in the compressed zone is considered as linear. In the context of this research, the master curves were identified from a parametric study, where the range of para-metric values concerning the tensile properties was focused

Fig. 4—Derivation of average and characteristic bilinear curves: proposed point-by-point method.


on common values for UHPFRC to be more precise—the tensile loss of linearity and ultimate tensile strength ranging from 6 to 14 MPa (870 to 2030 psi), the modulus of elasticity ranging from 50 to 65 GPa (7252 to 9427 ksi), and the tensile strain capacity ranging from 0.0005 to 0.0140. For each test configuration, 126 cases were investigated in the parametric study within the considered range of material parameters. Eighteen linear curves were obtained and used to plot the master curves. The MOR/ste-versus-strain-capacity master curve was also constructed with the same range of para-metric values concerning the tensile properties focused on common values for UHPFRC, as described previously. For each specimen size, 90 cases were investigated in the para-metric study, and 10 linear curves were obtained and used to plot the master curves. For each specimen group and for the average and characteristic curves, only the mean master curve was used to quantify the deviation in terms of strength and strain capacity between the simplified inverse analysis and the point-by-point method.

Fig. 5—Outline of Qian and Li (2008) method.

The final strain eend-simp of the average curves obtained with the Qian and Li (2008) simplified inverse analysis is, for each specimen group, the average of all eSpecimen-k-simp, where eSpecimen-k-simp is the strain identified with the mean master curve of tensile strain capacity versus deflection capacity for the specimen k.

The characteristic tensile stress-strain curve is obtained by applying the Qian and Li (2008) simplified inverse analysis to the characteristic bending-moment-versus-midspan-deflection curve with a final strain emin-simp equal to the minimum of all eSpecimen-k-simp, where eSpecimen-k-simp is the strain identified with the mean master curve tensile strain capacity versus deflection capacity for the specimen k.

In Tables 4 and 5, the average and characteristic bilinear tensile stress-strain relationships obtained from the proposed point-by-point inverse analysis and the Qian and Li (2008) simplified inverse method are presented for each specimen group. In considering the average stress of the postcracking part of the bilinear tensile stress-strain curve ((s1 + s2)/2), the stress overestimation induced by the mechanical assumption to convert the deflection into curvature and by the assumption of uniform stress distribution along the tensile height is, on average for all specimen groups, equal to 7% (with a maximum close to 10%) for the average curves and 10% (with a maximum close to 14%) for the character-istic curves. In terms of strains, the underestimation of the considered strain at crack localization obtained by the Qian and Li (2008) simplified inverse analysis procedure is, on average, equal to 15% (with a maximum close to 24%) for the average curves and 22% (with a maximum close to 40%) for the characteristic curves.

COMPARISON WITH DIRECT TENSILE TEST (DTT) RESULTS

A joint research program was recently completed by the U.S. Federal Highway Administration and the French IFFSTAR (formerly LCPC) to develop a DTT applicable to UHPFRC that covers the full range of uniaxial tensile behavior through strain localization and can be completed on cast or extracted specimens (Graybeal et al. 2012). In the context of this study, this DTT method was applied for all specimen groups, except for Specimen B2A-L. In Fig. 6, the average tensile stress-strain relationships obtained from the proposed point-by-point inverse method, the Qian and Li

Table 4—Average bilinear tensile stress-strain relationships (for each specimen group) derived from flexural tests associated with inverse analysis and obtained with DTTs

B2A-S B2A-L F1A-L F2A-L F1C-S F1C-L

Average curves

Proposed point-by-point method

sa1, MPa (ksi) 9.22 (1.33) 9.60 (1.39) 10.17 (1.48) 8.79 (1.28) 11.12 (1.61) 10.35 (1.51)

sa2, MPa (ksi) 11.36 (1.65) 11.87 (1.73) 10.59 (1.54) 9.18 (1.33) 11.13 (1.61) 11.31 (1.64)

e1 0.000153 0.000159 0.000184 0.000160 0.000203 0.000192

eend-ppt 0.008200 0.007400 0.008000 0.005400 0.006500 0.007600

Qian and Li (2008) simplified method

sa1, MPa (ksi) 11.00 (1.60) 11.12 (1.61) 10.99 (1.59) 9.92 (1.44) 11.99 (1.74) 11.91 (1.73)

e1 0.000185 0.000185 0.000203 0.000181 0.000233 0.000226

eend-simp 0.007100 0.006600 0.006100 0.004300 0.006100 0.006300

DTTs

sa1, MPa (ksi) 8.86 (1.28) — 10.00 (1.45) 8.60 (1.25) 10.30 (1.49) 10.50 (1.52)

sa2, MPa (ksi) 10.14 (1.47) — 10.00 (1.45) 9.00 (1.31) 11.10 (1.61) 11.55 (1.67)

e1 0.000144 — 0.000180 0.000155 0.000185 0.000194

eend-dtt 0.007400 — 0.004200 0.003000 0.004800 0.006000


increasing strain is predominantly due to increasing crack openings as opposed to further crack initiation.

To compare the results derived from the proposed inverse analysis method and the Qian and Li (2008) simplified proce-dure with the DTT results, bilinear curves were constructed from the sampled tensile stress-strain responses obtained with DTTs (refer to Tables 4 through 6).

(2008) simplified inverse analysis, and the average experi-mental curves obtained from the DTT are presented for each specimen group. During DTTs of Specimens B2A-S and F1C-L, after an early phase during which multi-cracking occurred, a hardening phase followed without new cracks. Indeed, as explained in Fischer and Li (2007), multiple crack formation can appear as stabilized when all available matrix flaws at ambient stress have been activated. Afterward, the

Fig. 6—Average tensile stress-strain curves for each specimen group: point-by-point inverse method, Qian and Li (2008) analysis, and DTT.

Table 5—Characteristic bilinear tensile stress-strain relationships (for each specimen group) derived from flexural tests associated with inverse analysis and obtained with DTTs

B2A-S B2A-L F1A-L F2A-L F1C-S F1C-L

Characteristic curves


sc1, MPa (ksi) 6.27 (0.91) 7.57 (1.10) 7.97 (1.16) 5.92 (0.86) 8.43 (1.22) 8.65 (1.26)

sc2, MPa (ksi) 6.27 (0.91) 8.66 (1.26) 8.59 (1.25) 8.27 (1.20) 8.46 (1.23) 10.55 (1.52)

e1 0.000104 0.000126 0.000144 0.000108 0.000154 0.000160

emin-ppt 0.007400 0.006800 0.006440 0.002840 0.004150 0.005150


sc1, MPa (ksi) 6.87 (1.00) 9.12 (1.32) 9.07 (1.32) 8.25 (1.19) 9.32 (1.35) 10.53 (1.52)

e1 0.000115 0.000152 0.000168 0.000151 0.000181 0.000200

emin-simp 0.004400 0.005600 0.004600 0.002500 0.003800 0.003900

DTTs

sc1, MPa (ksi) 7.50 (1.09) — 7.70 (1.12) 6.60 (0.96) 8.80 (1.28) 9.50 (1.38)

sc2, MPa (ksi) 7.82 (1.13) — 7.70 (1.12) 7.70 (1.12) 10.20 (1.48) 9.50 (1.38)

e1 0.000130 — 0.000140 0.000110 0.000160 0.000165

emin-dtt 0.005600 — 0.003000 0.001500 0.003900 0.005200


Table 6—General comparison of results derived from both inverse analysis methods with DTT results

Strength Pseudo-strain-hardening

Average curves


Global trend Overestimation Overestimation

Average deviation, % +3 +30

Maximum deviation, % +8 +48

Minimum deviation, % –2 +9


Global trend Overestimation Overestimation



Minimum deviation, % +7 –4.5

Characteristic curves


Global trend Underestimation Overestimation

Average deviation, % –5 +26


Minimum deviation, % –22 –1


Global trend Overestimation No trend



Minimum deviation, % –11 –35

In terms of strength, the proposed point-by-point inverse analysis method slightly overestimates the strength when considering average curves and underestimates the post-cracking stress when considering characteristic curves. The Qian and Li (2008) simplified inverse procedure slightly overestimates the stress for average and characteristic curves. In terms of strain, the Qian and Li (2008) simplified inverse method results are closer to DTT results than the proposed inverse procedure. Nevertheless, this smaller deviation is due to the co-existence of two opposed effects when considering the Qian and Li (2008) simplified inverse procedure:• The flexural tests involve an overestimation of the strain

capacity due to the fact that the side under higher tension corresponds to the zone where the preferential orienta-tion of fibers is optimal. This phenomenon has already been observed by Tailhan et al. (2004) on a multi-scale cement-based composite (MSCC). Completing the tests on larger prisms would minimize the strain gradient effect and thus would allow the results to be closer to the DTT results. Investigating this size effect was outside of the scope of this experimental program.

• As explained previously, the Qian and Li (2008) simpli-fied inverse method underestimates the real strain capacity in flexural configuration due to the mechanical assumption used to convert the deflection into curvature.

In terms of strength, the comparison between DTTs and both inverse methods based on flexural tests presents different results when considering the average or character-istic curves. This change is due to a “statistical size effect.” For the flexural tests, the tensile area is smaller than in the DTTs. As a consequence, on average, the results are better for flexural tests, but the impact of an eventual composite (matrix and fibers) flaw is greater and the standard deviation is more important. Thus, the characteristic strength can be inferior for flexural tests.

CONCLUSIONSThe research described herein presents a new method

based on flexural tests to determine the tensile stress-strain response of UHPFRC with pseudo-strain-hardening behavior in tension. This method uses two staggered LVDTs employed as extensometers to obtain the experimental bending-moment-versus-midspan strain at the tensile face curve and also to capture the crack localization. It is then associated with a point-by-point inverse analysis. Thus, the UHPFRC tensile stress-strain relationship is derived while minimizing the assumptions that can introduce deviations or artifacts in the results. In particular, assumption of the profile of the tensile stress-strain relationship is not required.

A comparison of this method with the Qian and Li (2008) simplified inverse procedure was completed on diverse UHPFRC specimens with different steel-fiber ratios or different curing regimes. The results show a slight strength overestimation and a strain underestimation during the pseudo-hardening phase (in flexural configuration) in this simplified inverse method, induced by the assumption of uniform stress distribution along the tensile height and by the mechanical assumption used to convert the deflection into curvature.

From the comparison with the companion DTT results, the following specific conclusions can be drawn:• The average tensile stress-strain response of UHPFRC

derived from flexural tests is slightly higher in terms of strength and strain capacity when compared with curves obtained from DTTs. This response results from a smaller tested tensile zone for flexure tests. Coincidently, the characteristic values can be inferior for flexural tests due to a larger standard deviation. This conclusion was demonstrated through the use of an inverse analysis method that minimizes the assump-tions, thus providing more realistic and reliable results.

• Using larger cross-section prisms would minimize the strain gradient effect and thus would likely facilitate greater coincidence in the accuracy of the flexure test


results as compared to the DTT results to provide reli-able design figures. In the case of thin elements made of UHPFRC under predominant direct tension and whose characterization of tensile postcracking behavior is real-ized from four-point flexural tests, the strain capacity has to be reduced to take into account the strain overes-timation (before reaching crack localization) due to the flexural test configuration.

The proposed method confirms the efficiency of an inverse analysis integrating direct strain measurements along with the possibility of defining a design stress-strain law within the appropriate validity range.

ACKNOWLEDGMENTSThis work was supported by the IFSTTAR and FHWA departments in

charge of Scientific Issues and International Relationships. The authors would thus like to acknowledge the support of H. Van Damme, S. Proeschel, P. Malléjacq (IFSTTAR), D. Elston, I. Saunders, and C. Richter (FHWA). They are also pleased to thank the teams from the IFSTTAR Structures Laboratory and the TFHRC Structures Laboratory for their technical help.

The publication of this article does not necessarily indicate approval or endorsement of the findings, opinions, conclusions, or recommendations either inferred or specifically expressed herein by FHWA, the U.S. govern-ment, IFSTTAR, or the French government.

NOTATIONb = prism widthE = elastic modulush = prism heightMc = bending moment of zone under

compressionMmax = maximum bending momentMt = bending moment of zone under tensionMt,i + 1, an,i + 1, fi + 1, etf,i + 1, = refers to loading step i + 1Nc = axial force of zone under compressionNc,i, Mc,i, Nt,i, Mt,i, an,i, fi, etf,i, = refers to loading step iNc,i + 1, Mc,i + 1, Nt,i + 1, = refers to loading step i + 1Nt = axial force of zone under tensionz = ordinate (vertical axis)an = relative height of zone under tensionet(z) = tensile strain at ordinate zetf = strain at extreme tension fiberetf – measured = average of two extensometers’

measurementf = curvaturesc(z) = compressive stress at ordinate zst(et) = stress corresponding to strain et

ste = effective tensile strength

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Title no. 109-M54


ACI Materials Journal, V. 109, No. 5, September-October 2012.MS No. M-2011-228 received July 19, 2011, and reviewed under Institute publication

policies. Copyright © 2012, American Concrete Institute. All rights reserved, including the making of copies unless permission is obtained from the copyright proprietors. Pertinent discussion including author’s closure, if any, will be published in the July-August 2013 ACI Materials Journal if the discussion is received by April 1, 2013.

A First-Cut Field Method to Evaluate Limestone Aggregate Durabilityby Julienne Ruth Emry, Robert H. Goldstein, and Evan K. Franseen

Crumpton 1981; Clowers 1999). D-cracking is character-ized by fine, closely spaced cracks parallel to joints at the surface of the concrete and is believed to be caused when porous aggregates in the concrete become saturated with water and are subjected to repeated cycles of freezing and thawing (Clowers 1999; Lamont and Pielert 2006; Glass 1990). The result of these studies was the establishment of a series of standard physical tests to determine aggregate durability (Wallace and Hamilton 1982) that were patterned after ASTM standard physical tests and published studies on aggregate durability (Clowers 1999). The implementation of these standard tests in Kansas led to a decrease in the propor-tion of highways with D-cracking, from 48% of those paved from 1961-1974 to currently less than 1% (Clowers 1999).

Other state, national, and international governmental agencies require similarly extensive testing to evaluate the physical properties of aggregate in response to long-term weathering (Lamont and Pielert 2006; Won 2005; EFNARC 2002; Koubaa and Snyder 1996). Many of the standard physical tests, such as ASTM C666/C666M-03 (2008), require the production of test concrete cylinders that are subjected to many freezing-and-thawing cycles in the lab (Lamont and Pielert 2006). Acquiring results from tests such as ASTM C666/C666M can take upward of 6 months due to both the time needed to produce and test the cylinders and the fact that labs often have backlogs of samples waiting for testing. These long wait times can lead to inadequate sampling of aggregates, which can lead to a failure to detect the great lateral and vertical variability that is typical of sedimentary carbonates (for example, McKirahan et al. [2003], Enos [1983], Heckel [2002], Moore [1935], Crowley [1969], and Franseen and Gold-stein [2004]). KDOT researchers have documented this high degree of variability within individual ledges (Cowers 1999; Wallace and Hamilton 1982) and have noted that, because of this variability, the samples used for physical tests may not be representative of all the material quarried.

Quality control issues and the costs associated with using low-quality aggregate are factors illustrating a growing need for effective first-cut field techniques to evaluate limestone aggregate durability. While KDOT requires that geolo-gists inspect aggregate-producing ledges once every 2 years (Clowers 1999), most quarries do not employ geologists to monitor aggregate quality during quarrying. Therefore, devel-oping field-based methods to determine aggregate durability

The demand for durable limestone aggregate and concerns about environmental sustainability are current industry issues. Lime-stone aggregate abundance, lithologic variability, and extensive testing by the Kansas Department of Transportation (KDOT) make Kansas an excellent locality for developing a field-based technique for assessing aggregate durability. This study documents a first-cut method for evaluating aggregate resistance to freezing and thawing prior to subjecting samples to time-consuming physical tests such as ASTM C666/C666M. Gamma-ray-spectrometry-measured potassium (K) radioisotopes on a quarry face were statistically determined to be predictive of aggregate freezing-and-thawing resistance. A logistic model based on maximum potassium value (Kmax) provided the best prediction of resistance to freezing and thawing, as described by the statistical likelihood that a lime-stone bed with a micritic matrix will pass or fail KDOT physical tests (KTMR-21 and ASTM C666/C666M). In areas of limestone production, where resistance to freezing and thawing is a concern, this fast, inexpensive, first-cut methodology could be calibrated to region-specific physical tests.

Keywords: D-cracking; durability; freezing-and-thawing resistance; gamma-ray spectrometry; sustainability.

INTRODUCTIONIndustrial demand for durable limestone aggregate for

state, county, and municipal projects is increasing in the United States. In the state of Kansas, limestone aggregate is an abundant resource that plays a significant role in the state’s economy. In 2008, almost 22 million tons (20 million tonnes) of crushed limestone valued at $171 million dollars was used or sold by Kansas aggregate producers (http://www.fhwa.dot.gov/engineering/geotech/hazards/mine/workshops/kdot/kansas01.cfm#table1). Although limestone aggregate production is an important industry in Kansas, concerns about limestone aggregate quality have led some municipalities to legislate the use of hard rock aggregate (for example, granite, rhyolite, and quartzite) imported from other states (http://www.kcmmb.org/Specs/specs.asp). Importing aggregate from other states when in-state aggre-gate resources are abundant and easily quarried can be costly and is a waste of local resources. In a review of the environ-mental impacts of concrete production, Mehta (2001) cited transportation as an environmental cost due to the heavy use of fossil fuels. Therefore, the use of local aggregate is desir-able for both environmental and economic sustainability.

Identifying and developing tools to predict the vari-ables that affect carbonate aggregate quality is an industry priority (Keyser et al. 1984; Klieger et al. 1974; Koubaa and Snyder 1996; Shakoor 1982). For example, persistent problems with D-cracking in concrete pavements prompted the Kansas Department of Transportation (KDOT) to conduct six studies between 1920 and 1980 to understand the factors involved in identifying and producing durable limestone aggregate (Bukovatz et al. 1973; Bukovatz and


clay minerals (illite, kaolinite, and smectite)—may reduce the freezing-and-thawing resistance of limestone aggregate.

McKirahan et al. (2000) suggested that a hand-held gamma-ray spectrometer could be used as a first-cut tool for evaluating aggregate durability, both due to its response to clay content in limestones and because measurements can be made rapidly on the face of the quarried ledge. A gamma-ray spectrometer measures the amount of the three major sources of natural gamma radiation in rocks (potassium, uranium, and thorium) along with the total gamma radiation. This device works on the principle of passive detection of the products of the radioactive decay series of 238U, 232Th, and 40K (Durrance 1986). The spectrometer used in this study (Fig. 1) uses an NaI crystal detector and a 137Cs refer-ence source with a precision of 0.01 nGyn/Hz to quantify the amount of the daughter products detected. The justification behind using this tool is based on three observations related to clay properties: 1) clay minerals have significantly higher potassium content than carbonates; 2) clay minerals are often associated with organic material that fixes uranium; and 3) some clay minerals can adsorb thorium (Doveton 1994).

Almost all methods for determining aggregate durability are lab-based, such as electrical resistivity (Sengul and Gjørv 2008), vacuum absorption (Williamson et al. 2007), and thermogravimetric methods (Dubberke and Marks 1994). Field-based methods are virtually nonexistent; there-fore, the purpose of this project was to evaluate the validity of using a gamma-ray spectrometer as a first-cut tool to eval-uate limestone aggregate freezing-and-thawing resistance. Spectrometer-derived measurements were compared to tests calibrated to a modified ASTM C666/C666M test run by KDOT. These data were used to develop a probability-based model to evaluate limestone aggregate quality.

RESEARCH SIGNIFICANCEEnvironmental and sustainability issues associated with

the increasing demand for limestone, dolomite, and marble (carbonate) aggregate, quality control issues, and the increased cost associated with using low-quality aggregate are all factors that illustrate a growing need for effective first-cut techniques that can predict long-term resistance to freezing and thawing. Such techniques can be used as a screening tool for standardized physical tests, such as ASTM C666/C666M, that are commonly used to evaluate aggregate freezing-and-thawing resistance. Gamma-ray spectrometry provides a successful probability-based first-cut field tool to evaluate limestone aggregate freezing-and-thawing resistance.

EXPERIMENTAL PROCEDURELocations and stratigraphic units tested

The selection of sites for this study was based on three primary criteria: location, accessibility, and whether the primary physical tests for rating aggregate quality (KTMR-21 test [KDOT 2007] and ASTM C666/C666M Test Procedure B) had been performed on the ledges. The 12 selected quarries are in eastern Kansas and western Missouri and are operated by seven different companies (Fig. 2). Stratigraphic intervals (ledges) from 10 different stratigraphic units were sampled to incorporate as much lithologic variability as possible.

Physical test protocolThe physical tests were run by KDOT on samples collected

by KDOT geologists. Samples were gathered to represent all

Julienne Ruth Emry is a PhD Student at the University of Kansas, Lawrence, KS. She received her BA from Colorado College, Colorado Springs, CO, in 1999, and her MS from the University of Kansas in 2006. Her research interests include field-based methods for predicting aggregate quality in carbonate aggregates, carbonate deposi-tional environments and sequence stratigraphy, and the formation and alteration of Archean rocks.

Robert H. Goldstein is the Haas Distinguished Professor in the Department of Geology at the University of Kansas. He received his BS in geology from Juniata College, Huntingdon, PA, in 1979, and his MS and PhD from University of Wisconsin-Madison, Madison, WI, in 1981 and 1986, respectively. His research interests include the stratigraphy and diagenesis of limestone and dolomite, and developing new tech-niques for the application of fluid inclusions in sedimentary rocks.

Evan K. Franseen is a Professor in the Department of Geology and Senior Scientific Fellow at the Kansas Geological Survey (University of Kansas). He received his PhD in geology from the University of Wisconsin-Madison in 1989. His research interests include sedimentology, stratigraphy, and diagenesis of carbonate rocks, and geologic applications for understanding aggregate durability.

that are both quantitative and easy-to-use is important to ensure long-term production of consistent aggregate products.

An initial project by the authors’ research team (McKi-rahan et al. 2000) focused on evaluating the factors that affect the quality of limestone aggregate. This study showed that all limestone textural classifications (Dunham 1962) may produce aggregate with high resistance to freezing and thawing. The presence of abundant fine-grained carbonate matrix (micrite), medium-grained carbonate matrix (microspar), or abundant spar (coarsely crystalline chemical precipitate) has no apparent impact on freezing-and-thawing resistance. Bulk spar percentage, spar size, insoluble residue percentage, and grain size have some control over freezing-and-thawing resistance, but are not reliable indicators of aggregate performance on the modi-fied ASTM C666/C666M freezing-and-thawing resistance test. Of all the variables examined, the strongest correlations between rock properties and modified ASTM C666/C666M physical test results were total clay percentage, clay distribu-tion, and composition of insoluble residues. Specifically, the higher the percentage of clay-rich limestone strata observed in the KDOT limestone bed, the poorer the performance in the standard KDOT physical tests. Exceptions to this trend were clays concentrated in stylocumulates (seams created by pressure dissolution) or clays in shaley beds, which may or may not have a negative impact on aggregate durability. McKirahan et al. (2000) hypothesized that the composition of insoluble residues—specifically, the presence of three

Fig. 1—Photo of handheld gamma-ray spectrometer used in this study. This specific model has channels to detect total gamma radiation, K, U, and Th.


strata in vertical intervals of rock ledges in quarries. KDOT breaks quarry ledges into KDOT beds based on parameters such as the general physical appearance of the rocks, basic rock type, and natural breaks in the ledges that affect produc-tion. These are essentially stratigraphic units that a quarry operator can recognize in a quarry face. KDOT beds vary in thickness, but are often essentially the same as the lithologic or stratigraphic beds that make up the limestone deposit. Multiple limestone layers, which can differ in lithology, can be combined into a single KDOT bed due to production issues, such as blasting techniques and earth-moving equip-ment capabilities. Physical tests are run on splits of the bulk samples taken from each KDOT bed.

KDOT uses the KTMR-21 test and a modified version of the ASTM C666/C666M Standard Test Procedure B to char-acterize the freezing-and-thawing resistance of aggregate resources. Aggregate samples rated as Class 1 are required for highway construction and must have a modified freezing-and-thawing value of at least 0.85 from the KTMR-21 test and durability factor of 95% or more, and an expansion value of 0.025 or less from the ASTM C666/C666M test. Current KDOT practice first involves testing the aggregate samples according to KTMR-21. The KTMR-21 Modified Freeze-Thaw Test is based on the method proposed by Scholer and Stoddard (1932). It involves sieving the aggregate to estab-lished specifications, weighing the aggregate, immersing it in water, and running the samples through 25 freezing-and-thawing cycles, after which the material is sieved and weighed again. The ratio of remaining weight of the sample to the original weight of the material is calculated to deter-mine the percent weight of the sample remaining after the test. For details of KTMR-21, “Soundness and Modified Soundness of Aggregates by Freezing and Thawing,” refer to Clowers (1999).

If the aggregate sample “soundness and modified sound-ness of aggregates by freezing and thawing” value is 0.85 or greater (85% of the sample weight is retained), then the sample is tested further, using modified ASTM C666/C666M Procedure B (Clowers 1999). Other test protocols, including specific gravity, adsorption, acid-insoluble residue, and Los Angeles abrasion are then often performed, but not required. The modified ASTM C666/C666M Procedure B test involves constructing test concrete cylinders from the aggre-gate samples and subjecting the cylinders to 300 freezing-and-thawing cycles, after which the durability factor and the expansion value (which is the percent length change of the cylinder) are calculated. KDOT designates rock classes based on the results of three of the physical tests performed on the samples.

Gamma-ray spectrometryQuarries commonly remove and stockpile shales that are

interbedded with limestones to access the limestone units for quarrying. During this process, shale particles become airborne and collect on the limestone ledges. Shale units, particularly the black shales in the area, are more radioac-tive than other rock types (Doveton 1994) and shale dust is common on the surfaces of the ledge. Most quarries also actively crush limestone into aggregate-sized pieces on site, creating a large amount of limestone dust that collects on the ledges, which could also affect the gamma-ray spectrom-eter readings. Therefore, to remove dust from each ledge, a cleaning procedure was employed before performing spec-trometry. The cleaning procedure used a 2700 psi (186.2 bars)

gasoline-powered power-washer supplied with water from a 50 gal. (189.3 L) pressurized cement-mixer tank. The ledges were power-washed for a maximum of 5 minutes until most of the water running down the ledge was clear.

After washing, 11.8 in. (300 mm) intervals were marked from the base to the top of the ledge to match the sampling radius of the spectrometer, thereby ensuring a continuous measurement of the natural gamma radiation of the rock (Geophysical Gamma-ray Spectrometer 2001). To avoid bias due to weathered material and material washed off the outcrop, the first measurement at the base of each outcrop and any readings of weathered material at the top of the ledges were discarded from the data sets.

A collection time of 3 minutes was used for each sample point to obtain an appropriate statistical sample for rocks rela-tively low in radiation (Geophysical Gamma-ray Spectrom-eter 2001). Uneven surfaces can affect the signal acquired by the spectrometer, so it was held stable on the surface of the rock at a relatively flat point on the outcrop face, as per the manufacturer’s recommendation (Geophysical Gamma-ray Spectrometer 2001). The sampling radius of the instrument used was 11.8 in. (300 mm) and this was used to determine the sample spacing on the outcrop (Geophysical Gamma-ray Spectrometer 2001). Data generated by the gamma-ray spectrometer included total gamma radiation, potassium, uranium, and thorium in both counts-per-second (nGyn/Hz), and concentration (ppm or %) for each of the 948 sample points in 22 stratigraphic sections. All spectrometry data

Fig. 2—Index map of quarry locations in Kansas and Missouri. Multiple ledges were sampled at most locations.


Stratigraphic descriptionDetailed measurements and descriptions of the rock ledges

were recorded in the field and reinforced with thin-section petrography. Six rock-type classes were differentiated on the basis of clay content, distribution of clay, and whether the material between depositional grains in the limestone was micrite (Dunham 1962) or sparry calcite. Micrite consists of tiny crystals of calcite that are not discernible with a 10× hand lens. It appears as opaque, solid-looking mate-rial surrounding coarser depositional grains in a limestone. Sparry calcite consists of more translucent, coarsely crys-talline calcite in which individual crystals are visible with a 10× hand lens (Fig. 3). This study analyzed only lime-stone materials that contained micritic matrix as a means of limiting the number of variables.

Clays are typically distributed in three different ways in limestones: disseminated clay, diffuse stylocumulates, and concentrated stylocumulates or shale beds (Fig. 4). Stylo-cumulates are clay-rich zones that occur within limestone beds, commonly along bedding planes created by the pres-sure dissolution processes. The six rock-type classes defined in this study are limestones with 1) micrite matrix, dissemi-nated clays, and diffuse stylocumulates; 2) micrite matrix with disseminated clays; 3) micrite matrix only; 4) micrite matrix with diffuse stylocumulates; 5) sparry calcite matrix, which is disseminated clay-poor or diffuse stylocumulate-poor; and 6) shale and siltstone.

Statistical analysisLinear regression analysis was used to test for a relationship

between spectrometer measurements and the various KDOT test measurements (KTMR-21 and modified ASTM C666/C666M). Logistic analysis was used to test for a relationship between spectrometer measurements and a KDOT bed’s pass-or-fail status.

RESULTS AND DISCUSSIONSpectrometry—Kmax model

A total of 948 spectrometer readings were taken of potas-sium (K), uranium (U), thorium (Th), and total radiation. Both the median and maximum value for K, U, Th, and total radiation for each KDOT bed were used to test for a relation-ship between spectrometry measurements and the modified ASTM C666/C666M and KTMR-21 test results. Spectrom-etry measurements tended to show higher and more vari-able values where strata failed the KDOT physical tests and where strata contained the highest stratigraphic thicknesses of clay-rich strata. This result can be illustrated by comparing stratigraphic section descriptions from a location where samples passed the KDOT tests to achieve Class 1 designa-tion (Fig. 5(a)) to a stratigraphic section description from a location where samples failed the KDOT tests (Fig. 5(b)). Linear regressions were performed on summary statistics (maximum and median) to determine the degree to which spectrometry values were related to the ASTM C666/C666M, KTMR-21, and acid-insoluble residue values. Only 10 out of 48 linear regressions were statistically significant at the 0.05 confidence level between physical tests and spectrom-eter measurements. The r2 values for all of the significant tests were weak (maximum r2 value of 0.4536), indicating that spectrometer measurements do not predict KDOT phys-ical test results accurately with a linear model. This result is not surprising considering that a single, thin, low-durability stratigraphic interval within a KDOT bed may lead to aggre-

are presented in the Appendix* and are organized by quarry location and KDOT bed numbers.

Summary statistics of the gamma radiation data from each KDOT bed were used for statistical comparison to the KDOT physical test results. Included in the Appendix are the KDOT physical test pass/fail status (where “pass” designates samples that passed all the requirements for Class 1 designa-tion and “fail” designates samples that did not meet one or more of the aforementioned requirements); the three main KDOT test result values (freezing-and-thawing soundness, expansion, and durability factor); and the summary statis-tics for each KDOT bed. At locations where it was possible to measure centimeter-resolution stratigraphic sections, the percent of clay-rich rock is also included.

* The Appendix is available at www.concrete.org in PDF format as an addendum to the published paper. It is also available in hard copy from ACI headquarters for a fee equal to the cost of reproduction plus handling at the time of the request.

Fig. 3—Example of sparry calcite (A) and micritic matrix (B) facies. Photo scale and detail are consistent with 10× hand lens. Sparry matrix commonly looks glassy, is often clear, and will reflect points of light. Micritic matrix appears opaque and individual grains and crystals cannot be seen with 10× hand lens. (Note: M is micritic matrix; S is sparry matrix; G is grain; and Rf is point of reflected light.)

Fig. 4—Hypothetical illustration of two limestone beds with various forms of clay distributed within them: (1) concen-trated stylocumulates or thin shale beds are typically located along bedding planes and may branch into surrounding limestones; (2) concentrated stylocumulates can also occur within limestones; (3) these commonly branch into slightly more diffuse stylocumulates near their ends (3a) or have zones of diffuse stylocumulates within them (3b); and (4) diffuse stylocumulates also occur as thin wisps or stringers of clay-rich material within limestones and may have “horse-tail” appearance (4a) (after McKirahan et al. [2000]).


gate failure, and that overcoming a compositional or mechan-ical threshold rather than linear degradation in rock properties may lead to failure. Thus, a different approach to statistical analyses was needed to determine if spectrometry measure-ments could reliably indicate if a limestone bed would pass or fail the physical tests.

As gamma radiation measurements are continuous vari-ables and pass/fail status is a categorical variable, logistic regression was used to compare gamma radiation values to whether a limestone bed passed (1) or failed (0) the phys-ical tests for aggregate freezing-and-thawing resistance (Sokal and Rohlf 1995; Gotelli and Ellison 2004). Logistic regression analyses were performed on all of the individual summary statistics for measured K, U, Th, and total radia-tion, and on several combinations of them. The results indi-cate that maximum potassium value for a limestone bed (Kmax) is the most useful gamma radiation measurement and will be the focus of the remaining discussion (Fig. 6). All additional analyses are included in the Appendix.

Fig. 5—Measured sections of middle of Argentine limestone (A) in Hunt-Midwest quarry in Crawford, KS, and Ervine Creek limestone from Martin Marietta quarry in Big Springs, KS (B). Eight of the 10 symbols shown in symbol key (crinoids, phylloid algae, oncoids, small stromatolites, high-spired gastropods, brachiopods, coated shell fragments, and fenestrate bryozoans) represent fossil organisms that are commonly found in rocks of this geologic age. Stylocumulates are features formed by pressure dissolution processes that affect rocks during burial. Gray rectangle denotes less than 4 in. (100 mm) shale bed in Ervine Creek limestone. Vertical spacing between sample points is 11.8 in. (300 mm). Potassium values for KDOT beds in weight % are plotted increasing from left to right on x-axis. Kmax values in spectrometer data for KDOT beds are circled. Measured stratigraphic sections and values for physical tests for expansion and durability (modified ASTM C666/C666M) are shown to the left of the spectrometer values. KDOT Bed 5 in Argentine limestone (A) passed KDOT physical tests and upper part of KDOT Bed 6 was not tested at this locality. KDOT Bed 2 in Ervine Creek Limestone failed KDOT physical tests, and KDOT Bed 3 failed initial KTMR-21 protocol, so it was not tested for expansion and durability using modified ASTM C666/C666M test.

Fig. 6—Plot of Kmax versus Pass (y-coordinate of “1”) or Fail (y-coordinate of “0”) for all KDOT beds. Several ledges of each stratigraphic unit at different quarries were tested. Individual stratigraphic units are indicated by different shaped symbols. Note that anomalous reading for Upper Farley limestone at Hunt Midwest Sunflower quarry (UFarley_at_HMSun) was dropped from logistic model, as described in the text.


the highest value for Kmax (Fig. 6). This lack of unique ranges in Kmax values for individual stratigraphic units supports the general applicability of the proposed model.

The logistic model shows that the probability of attaining a Class 1 designation (passes the physical tests) decreases as the Kmax value increases (Fig. 7). The equation for this model is

2.81 ( 9.27)

2.81 ( 9.27)1

max

max

K

K

epe

+ −

+ −=+

(1)

where p is the probability that a limestone bed with a given Kmax value will pass the physical tests for Class 1 designa-tion. Logistic models allow the user to define upper and lower threshold values that provide the appropriate risk for decision-making purposes. Based on consultation with KDOT and quarry personnel, an 80% or greater probability of passing was chosen as a conservative threshold value. Thus, given the 80% threshold, a Kmax value in Region A (Fig. 7) will have an 80% or greater probability of passing the KDOT physical tests (or a 20% or less probability of failing). Any ledge with a Kmax value in Region C will have 80% or greater probability of failing the KDOT physical tests. Regions A and C (the top and the bottom of the curve) are characterized by relatively flat slopes, indicating consis-tently high or low probabilities of passing KDOT physical tests. In contrast, the steep slope in Region B indicates that a small change in Kmax would produce significant changes in the probability of a limestone bed passing or failing the KDOT physical tests. Using the 80% threshold, there are four KDOT beds that have Kmax values that fall within Region A (Fig. 7). Three out of these four KDOT beds passed the physical tests required for class one aggregate (x-coordinate of 1), which illustrates that at the 80% threshold, the model would have correctly predicted that a KDOT bed would pass or fail the physical tests 75% of the time. Applying the same 80% threshold value for Region C (a 20% probability or less that the bed would pass the KDOT physical tests, or an 80% probability or greater that they would fail), there are six KDOT beds with Kmax values that fall into this region (Fig. 7). Note that all six KDOT beds are shown to have failed the KDOT physical tests (x-coordinate of 0), which illustrates that the model prediction was correct 100% of the time. This is a critical result, as it illustrates the accuracy of the model for predicting low-quality aggregate, which could be very useful as a first-cut test to identify aggregate resources of exceedingly low quality. The remaining data lie in Region B (Fig. X) and had probability values that did not meet the 80% threshold value (either a probability of passing 80% of the time or failing 80% of the time). Thus, the tech-nique’s function as a first-cut technique is diminished with mid-range Kmax values. Using the arbitrary 80% threshold value, the technique does not perform well predicting the pass/fail status of mid-range Kmax values. It performs well with low Kmax values, which predict that resources will pass the physical tests; and performs extremely well with high Kmax values, which identify resources that will fail the physical tests.

A spreadsheet that uses Kmax to calculate the probability of passing or failing the KDOT physical tests of aggregate durability in limestone is available (Emry et al. 2006). Other

Out of the 948 measurements, one anomalously high Kmax value from the upper Farley Limestone at the Hunt Midwest Sunflower quarry was removed from the analyses (Gotelli and Ellison 2004; Sokal and Rohlf 1995). The upper Farley shows significant lateral variation within this quarry and the area from which this anomalously high Kmax value was generated is approximately 16.4 ft (5 m) from where the original samples were taken for the KDOT physical tests. It is therefore likely that the rock tested by KDOT 2 years prior was very different than the rock analyzed with gamma-ray spectrometery in this study.

Sampling 10 different geologic formations or members was done by design to demonstrate the broad applicability of the technique to limestones with micritic matrix, and to avoid concerns that the technique might only be applicable to a very specific group of limestones in a very specific geographic region. In the initial analysis, there was a concern that each of the 10 stratigraphic units evaluated might have their own unique relationship between Kmax values and aggre-gate durability, which could necessitate having a separate model for each rock ledge. To address this potential bias by stratigraphic unit, data for each stratigraphic unit were exam-ined separately. It is apparent that each formation includes a broad range of values, and no formation has values that are clumped within a small range (Fig. 6). For example, the Argentine Limestone includes one of the lowest values and

Fig. 7—Graph of logistic model of relationship between maximum value for potassium (Kmax) and data from which it was derived. Data plotted along y-coordinate at “0” are those samples that failed and those plotted along y-coordi-nate of “1” are samples that passed. Dashed lines dividing plot into regions (A, B, and C) represent arbitrary 80% threshold values that beds would pass or fail physical tests. Region A represents values that have ≥80% probability of passing physical tests. Region C represents values that have ≤20% probability of passing (or ≥80% probability of failing). Region B represents values that fall between ≥80% probability that bed will pass or fail physical tests. Steep slope in Region B indicates that small change in Kmax would produce significant changes in probability that limestone bed would pass or fail physical tests and illustrates that less confidence should be given to probability predictions in this range. Region B also represents that portion of model where prediction of whether bed will pass or fail physical tests at given threshold is not within that established threshold value. Any threshold value quarry operators or governmental offi-cials deem useful depending on factors not addressed in this study could be applied to this model.


threshold values could ultimately be used depending on cost-benefit factors not addressed in this study.

This method is valuable in that it saves time and resources by highlighting the best potential resources and, perhaps most importantly, clearly identifying the worst. It would be useful as a monitoring device to evaluate lateral changes in rock properties during quarrying operations, or could be useful in evaluating a new resource. Ultimately, it could help prevent the incorporation of low-quality limestone aggregate into portland cement concrete pavement. This method is not intended to replace physical tests, but by being able to use an appropriate threshold value, quarry operators would be able to high-grade samples likely to pass the physical tests and move them to the front of the line for production and testing. It also allows operators to identify rock that would have a high probability of failing the physical tests and either remove them from consideration for production or fast-track them for testing to rule them out as potential resources to exploit.

Visual inspection by a trained geologist to assess clay content is an alternative first-cut approach, but can be imprac-tical. Changes in disseminated clay may not be readily visible to a geologist without time-consuming laboratory analyses, and many quarries do not employ full-time geologists to visually inspect ledges during quarrying. Government geolo-gists may also have a long inspection rotation (for example, 2 years for KDOT geologists), which can add significantly to the time in which a questionable ledge must wait before it can be inspected and a decision made as to whether it should undergo physical testing. The spectrometry method provides a fast, inexpensive, quantitative, and reproducible approach to monitor limestone ledges during quarrying, and it can be performed on site in real time by existing personnel. Due to the fact that this method correlates spectrometry readings to pass/fail status instead of specific test protocols, it can be easily adapted for use in other states or countries that have different test standards for aggregates in portland-cement concrete.

APPLICATIONThere are a variety of scenarios in which the spectrom-

eter percent potassium (K%) values would be useful. For example, as a quarry development tool, K% data can be acquired as soon as a ledge is opened to obtain a base-line value for material to be correlated with physical tests. Subsequent spectrometer readings should be compared to the original values to track quality control of the ledge as it is quarried. Also, as a visual change is seen in a quarried ledge, K% measurements are warranted, enabling govern-ment personnel and quarry operators to test the probability that the ledge would continue to pass physical tests.

The spectrometry data could also be used to track local and regional consistency or discover inconsistencies in K% values within a ledge. This may help identify consistent sources of durable aggregate or at the very least identify nascent problems due to the natural geologic variability seen within limestone beds. Tracking consistency in this manner should help to ensure that inferior aggregate is not included in aggregates that are used for highway construction. By taking readings in new quarries on newly exposed ledges, the methodology would be useful as an aggregate resource exploration tool. For example, if the Kmax value predicts a probability of passing below the accepted threshold value, then resources could be directed to more likely candidates.

If the spectrometer methodology is broadly implemented, a gamma-ray measurement database for each stratigraphic unit

could be developed. Building such databases would help to refine the statistical model proposed herein, and could be used to produce maps related to resource quality. As predictive methods for identifying future limestone aggregate resources are virtually nonexistent, this quantitative method has poten-tial to be an important time- and resource-saving tool.

The application requires first power-washing a quarry face and identifying that the limestone has a micritic matrix. The spectrometer is then used to measure K% values. If Kmax values for a specific limestone ledge fall within Region A (Fig. 7), the ledge has a high probability of producing highly durable aggregate. Resources that fall into this category could then be high-graded for physical testing. If values are in Region C, the ledge has a high probability of producing poor aggregate subject to degradation during freezing-and-thawing conditions, and these ledges could be either ruled out entirely or high-graded in the testing queue to rule them out using physical testing. This could be particularly impor-tant, especially if the ledges in question would need to be removed to access geologically lower units that may produce higher-quality aggregate. If Kmax values are in Region B, then no decision should be made on the basis of the Kmax values and it would be up to quarry operators or governmental agencies to decide how and when physical tests should be performed on these ledges.

CONCLUSIONSBased on the findings of this study, the following conclu-

sions can be made:1. The study’s results suggest that Kmax values obtained from

gamma-ray spectrometry can be used to determine if a given limestone bed with micritic matrix will pass or fail standard physical tests of aggregate freezing-and-thawing resistance.

2. Kmax values appear to correlate with clay content, which is a primary factor in the freezing-and-thawing resistance of limestone aggregate. Porosity in limestone aggregates, espe-cially those with sparitic matrix, can also be a factor in resis-tance to freezing and thawing. Focusing on limestones with a micritic matrix can help mitigate durability issues associ-ated with porous limestone aggregates.

3. This project established a novel, field-based method based on gamma-ray spectrometry to predict limestone aggregate freezing-and-thawing resistance. A predictive model using logistic regression of spectrometer data is viable for first-cut evaluation of the probability that a limestone bed will pass or fail freezing-and-thawing resistance tests. It is not intended to replace standard physical tests; instead, it maximizes efficiency in time and resources used to perform physical tests. This allows for faster characterization and identification of high-quality aggregate resources, and allows low-quality aggregate to be eliminated from production.

4. As logistic regression allowed for a threshold value for Kmax to be established in this study, it is reasonable to hypothesize that these values relate to threshold amounts of disseminated clay in limestones that lead to aggregate with high or low freezing-and-thawing resistance.

ACKNOWLEDGMENTSThis project was funded by KDOT Grant K-TRAN: KU-03-2. The

Kansas Geological Survey also provided support for this project during the first author’s employment as a visiting scientist. R. Henthorne provided logistical help, R. Houser provided GIS advice, J. Kelly provided statis-tical analysis and sampling methodology advice, D. Powell helped with X-ray diffraction, and G. Macpherson provided access to lab equipment. Thank you to J. Emry, who provided field assistance and useful comments on the manuscript. The authors also want to thank G. Lane; F. Rockers


from Shawnee Rock Co.; D. Maroney, R. Bryant, J. R. Downs, B. Foster, J. Kagarice, J. Nicholson, and C. Reed from Martin Marietta; T. Degonia, R. Stanley, and J. Epperson from Ashgrove Aggregates: the Stanley family for access to their private quarry, D. Patton from APAC; J. Crowley; J. Ciero at the Hunt Midwest Crawford Quarry; and R. Gonzales from Hamm Quar-ries for quarry access.

NOTATIONKmax = maximum potassium value in weight percent for KDOT bed measured with gamma-ray spectrometerK% = weight percent potassium measured with gamma-ray spectrometer

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Title no. 109-M55




Investigation of Properties of Engineered Cementitious Composites Incorporating High Volumes of Fly Ash and Metakaolinby E. Özbay, O. Karahan, M. Lachemi, K. M. A. Hossain, and C. Duran Atis

negative effects of higher cement content in ECC produc-tion, Yang et al.11 replaced cement with high volumes of fly ash (FA) (up to 85% by weight) and concluded that both the crack width and free drying shrinkage were reduced with increased FA content. Increasing the FA/PC ratio up to 5.6 decreased compressive strength (from 52.6 to 21.4 MPa [7.63 to 3.10 ksi]) and tensile strength (from 5.7 to 3.5 MPa [0.83 to 0.51 ksi]) and increased chloride-ion permea-bility12 drastically while reducing drying shrinkage to approximately 1000 × 10–6. However, these negative effects of high-volume FA on the mechanical- and durability-related properties of ECC may be remedied with the binary use of FA and metakaolin (MK).

MK is an ultra-fine material produced by the dehydroxyl-ation of a kaolin precursor by way of heating it to 650 to 900°C (1202 to 1652°F).13,14 MK is a silica-based product that, on reaction with Ca(OH)2, produces CSH gel at ambient temperatures. MK also contains alumina that reacts with CH to produce additional alumina-containing phases, including C4AH13, C2ASH8, and C3AH6. This pozzolanic material has been extensively investigated, particularly in relation to its effects on the durability and other properties of portland cement (PC) composites.15,16 Previous research has shown that the inclusion of MK in concrete remarkably improves early-age mechanical properties and enhances resistance to alkali-silica reaction,17 sulfate attack,18 and chloride-ion permeability.19 Introducing high-reactivity MK into concrete also ameliorates the energy absorption or toughness of high-performance steel fiber-reinforced concrete.20 Therefore, for applications where both enhanced durability and high toughness are required, the use of high-reactivity cementi-tious composites containing MK may be advantageous.13,20

This study investigated the binary uses of FA and MK in the production of ECC. ECC mixtures with two different FA + MK-PC ((FA + MK)/PC) ratios (1.2 and 2.2 by weight) were prepared by keeping the FA/MK ratio at 4.5. The investigation focused on the experimental characterization of compressive and flexural strengths, drying shrinkage, water absorption (WA), water porosity (WP), sorptivity, and chloride-ion permeability of ECC incorporating FA and MK. Two types of ECC mixtures—standard (FA/PC = 1.2)

This study was carried out to develop engineered cementitious composites (ECCs) incorporating binary blends of high volumes of fly ash (FA) and metakaolin (MK) for the purpose of achieving low drying shrinkage and high composite strength with adequate ductility and improved durability. ECC, an ultra-ductile cement-based composite reinforced with short random fibers, exhibits strain-hardening and multiple-cracking behavior in uniaxial tension and bending. Standard (M45) and high-volume FA ECC mixtures are typically produced by replacing portland cement (PC) with 55% and 70% of FA, respectively (FA-to-cement ratio of 1.2 and 2.2 by weight). In this study, the (FA + MK)/PC ratio was maintained at 1.2 and 2.2 and the FA/MK ratio was maintained at 4.5. Two replacement levels of MK with FA were adopted. The investigation used 10% and 12.5% MK by weight of total binder content, respectively. For the purposes of comparison, standard and high-volume FA ECCs were also studied. To determine the effect of binary blends of FA and MK on the properties of ECC, this study focused on the evaluation of free drying shrinkage, flexural and compressive strengths, porosity and water absorption (WA), sorptivity, and chloride-ion permeability. The experimental results showed that the drying shrinkage, porosity, absorption, sorp-tivity, and chloride-ion permeability properties were significantly reduced with the use of binary blends of FA and MK, while ECC’s ultra-high ductility and strain-hardening properties were preserved at an adequate level.

Keywords: drying shrinkage; durability; engineered cementitious composites; metakaolin.

INTRODUCTIONEngineered cementitious composites (ECCs) consti-

tute one of the most significant developments in the field of strain-hardening fiber-reinforced cementitious compos-ites and are microstructurally tailored based on the micro-mechanics design theory.1-4 Their strain-hardening and multiple-cracking behavior is characterized by a higher load-carrying capacity after first cracking of the matrix, which is associated with the appearance of closely spaced multiple cracks until composite peak load is reached.5 The tensile strain capacity of ECCs is 2 to 5%—several hundred times that of normal concrete. The compressive strength ranges from 50 to 80 MPa (7.25 to 11.6 ksi), depending on the composition of the mixture, which puts ECCs in the same class as high-strength concrete materials but without the associated brittleness.6 To obtain strain-hardening and multiple-cracking behavior, only a small amount of fine sand needs to be added to the matrix to control fracture tough-ness.7-10 Coarse aggregates are eliminated from the mixture, resulting in a higher cement content than that of conven-tional concrete. As a result of this special requirement, a high-drying shrinkage strain must be developed during the setting and hardening of the composite.7 To eliminate the


and high-volume FA (FA/PC = 2.2)—were also studied for comparison purposes.

RESEARCH SIGNIFICANCEECC is a newly developed, high-performance, fiber-

reinforced cementitious composite with substantial bene-fits in terms of improved ductility and durability due to its minimum crack width. To obtain strain-hardening and multiple-cracking behaviors, only a small amount of fine sand needs to be added to the matrix to control the frac-ture toughness. Coarse aggregates are eliminated from the mixture, resulting in higher cement content than conven-tional concrete. Therefore, a high drying shrinkage strain must be developed during setting and hardening of this unique composite. Researchers tried to decrease the drying shrinkage of ECC mixtures by using high volumes of FA; however, this resulted in a decrease in the mechanical and durability properties. This study focused on the binary use of FA and MK in ECC production. The negative effects of high-volume FA on the mechanical- and durability-related properties of ECC may be overcome with the binary use of FA and MK.

EXPERIMENTAL PROCEDUREMaterials and mixture proportions

Two groups of ECC mixtures were prepared. The first group included standard and high-volume FA ECC, which incorpo-rated Type I PC; Class F FA; normalweight microsilica sand (with an average and maximum grain size of 110 and 200 mm [0.004 and 0.008 in.], respectively); water; polyvinyl alcohol (PVA) fibers; and a polycarboxylic-ether-type high-range water-reducing admixture (HRWRA) with a solid content of approximately 30%. The second group of ECC mixtures was produced by replacing FA with MK at (FA + MK)/PC ratios of 1.2 and 2.2 while maintaining a constant FA/MK ratio of 4.5. The chemical composition and physical properties of the PC, FA, and MK used in this study are presented in Table 1. The mixture proportions of the first and second group of ECCs can be found in Table 2. All four mixtures contained 2% fiber content by volume. The fiber used in this study was an 8 mm (0.31 in.) long PVA fiber 39 mm (0.002 in.) in diameter with a tensile strength of 1600 MPa (235 ksi) and

Erdogan Özbay is an Associate Professor in the Civil Engineering Department at Mustafa Kemal University, Antakya, Turkey. His research interests include durability of concrete, use of waste materials in concrete, and self-consolidating concrete.

Okan Karahan is an Assistant Professor in the Civil Engineering Department at Erciyes University, Kayseri, Turkey. His research interests include construction materials and concrete technology.

ACI member Mohamed Lachemi is a Professor of Civil Engineering and Dean of the Faculty of Engineering and Architectural Science at Ryerson University, Toronto, ON, Canada. He is a member of ACI Committees 231, Properties of Concrete at Early Ages, and 237, Self-Consolidating Concrete. His research interests include the use of high-performance materials in the built infrastructure, including the development and use of self-consolidating concrete in construction.

ACI member Khandaker M. A. Hossain is an Associate Professor in the Department of Civil Engineering at Ryerson University. He is a member of ACI Committees 213, Lightweight Aggregate and Concrete, and 232, Fly Ash and Natural Pozzolans in Concrete. His research interests include sustainable construction, high-performance/self-consolidating concrete, reinforced concrete, and thin-walled composite structures.

Cengiz Duran Atis is a Professor of Civil Engineering and Dean of the Faculty of Engineering at Abdullah Gul University, Kayseri, Turkey. His research interests include construction materials and reinforced concrete structures.

Table 1—Characteristics of cement, FA, and MK

Chemical composition Cement FA MK

Sum (SiO2 + Al2O3 + Fe2O3) 27.60 85.60 95.00

SiO2, % 19.60 59.50 61 to 64

Al2O3, % 4.90 22.20 30 to 32

Fe2O3, % 3.10 3.90 1.10

CaO, % 61.40 5.57 0.40

MgO, % 3.00 — 0.30

SO3, % 3.60 0.19 0.05

Alkalis as Na2O, % 0.70 2.75 1.35

Loss on ignition, % 2.30 0.21 0.95

Physical properties

Blaine, cm2/g 3870 3060 13,900

+45 µm, % 3.00 9.60 1.20

Density, g/cm3 3.15 2.18 2.55

Notes: 1 cm2/g = 0.155 in.2/g; 1 mm = 0.0000393 in.; 1 g/cm3 = 168.45 lb/yd3.

Table 2—Mixture properties of ECC mixtures

Ingredients ECC-1_FA/PC = 1.2 ECC-2_FA/PC = 2.2 ECC-3_(FA + MK)/PC = 1.2 ECC-4_(FA + MK)/PC = 2.2

Water (W), kg/m3 331 327 326 318

Portland cement (PC), kg/m3 570 386 558 375

FA, kg/m3 684 847 547 673

MK, kg/m3 — — 122 150

Silica sand (S), kg/m3 455 448 446 435

Fiber (PVA), kg/m3 26 26 26 26

HRWRA, kg/m3 4.9 3.7 7.5 6.5

FA, % 55 69 45 56

MK, % — — 10 12.5

FA/PC 1.2 2.2 0.98 1.80

FA/MK — — 4.5 4.5

(FA + MK)/PC — — 1.2 2.2

Water-cementitious material ratio (w/cm) 0.27 0.27 0.27 0.27

Note: 1 kg/m3 = 1.6845 lb/yd3.


a density of 1300 kg/m3 (2190.5 lb/yd3). The fiber surface was coated with 1.2% oil by weight to reduce the fiber-matrix chemical and friction bond.21

The water-cementitious material ratio (w/cm) in all mixtures was controlled at 0.27. Slight adjustments in the amount of the HRWRA in each mixture were made to achieve consistent rheological properties for better fiber distribution and workability. Therefore, all specimens in the first and second groups of ECC had fresh properties similar to those in self-consolidating performance.11 As seen in Table 2, ECC mixtures incorporating FA and MK showed a higher HRWRA demand than those containing only FA. MK, a soft material made primarily of amorphous silicon dioxide and aluminum oxide and produced by the decom-position of kaolin at a temperature of 650 to 900°C (1202 to 1652°F), tends to absorb water to form kaolin. After the MK was added to the PC, it prompted cement hydration, which shortened the setting time. However, the MK-blended mixtures needed more water to achieve the same work-ability. Therefore, MK fluidity was degraded when the same dosage of HRWRA was added for the same workability; the study concluded that MK-blended mixtures require more HRWRA.22 Moreover, ECC mixtures with an FA/PC ratio of 1.2 had higher HRWRA demand than those with an FA/PC of 2.2.23 The smooth surface characteristics and spherical shape of the FA improved the workability characteristics of ECC mixtures so that similar workability properties at a constant w/cm were achieved by using a lower HRWRA content at a higher FA replacement level.24

A mortar mixer was used in the preparation of all ECC mixtures in this study. Solid ingredients, including cement, mineral admixture (FA or FA/MK), and aggregate, were initially mixed at 100 rpm for 1 minute. Water and HRWRA were then added into the dry mixture and mixed at 150 rpm for 1 minute and then mixed at 300 rpm for another 2 minutes to produce a consistent and uniform ECC matrix (without PVA fiber). Finally, PVA fiber was added and mixed at 150 rpm for an additional 3 minutes.

Specimen preparation and testingSeveral 285 x 25 x 25 mm (11.22 x 0.985 x 0.985 in.) bar

and 355 x 50 x 76 mm (13.97 x 1.97 x 2.99 in.) prism speci-mens from each mixture were prepared for drying shrinkage and four-point bending tests, respectively, and 100 x 200 mm (3.93 x 7.87 in.) cylinder specimens were prepared for rapid chloride permeability testing. Fifty mm (1.97 in.) cubic spec-imens were prepared to determine compressive strength, WA, water sorptivity, and WP. All specimens were demolded at the age of 24 hours and cured in sealed plastic bags at 95 ± 5% relative humidity (RH) and 23 ± 2°C (73°F ± 3.6°F) for 7 days. They were then air-cured at 50 ± 5% RH and 23 ± 2°C (73°F ± 3.6°F) for 28 days prior to testing. The complete testing program is detailed in the following sections.

Compressive and flexural strengthsThe compressive strength of the ECC mixtures was

determined by testing at least three 50 mm (1.97 in.) cubic specimens at the age of 28 days according to the procedure described in ASTM C39-94.25 A four-point bending test was performed under displacement control at a loading rate of 0.005 mm/s (0.0002 in./s) on a closed-loop controlled servo-hydraulic material test system. The span length of the flexural loading was 304.8 mm (12 in.) at the tension surface with a 101.6 mm (4 in.) center-span length at the compression

surface. During the flexural tests, load and midspan deflec-tion were recorded on a computerized data recording system.

Drying shrinkageDrying shrinkage measurements for all ECC mixtures

were made on three 285 x 25 x 25 mm (11.22 x 0.985 x 0.985 in.) bars up to 120 days after an initial curing of 1 day in the mold and 27 days in lime-saturated water in accor-dance with ASTM C157/C157M-04.26 The drying shrinkage specimens were stored in a drying room at 23 ± 2°C (73°F ± 3.6°F) and 50 ± 4% RH.

WA and porosityWA was determined as per ASTM C642-0627; speci-

mens were initially oven-dried at 105 ± 5°C (222°F ± 9°F) for 72 hours to reach constant mass and obtain oven-dry mass (W1). They were then immersed in water for 72 hours and the saturated surface-dry mass (W2) of the specimens was measured. The WA of each specimen was calculated as follows

2 1

1

W WWA (%) 100

W −

= × (1)

To determine the WP, the hydrostatic weight (W3) of the ECC specimens was also determined and the WP was calcu-lated as follows

2 1

2 3

W WWP (%) 100

W W −

= × − (2)

SorptivityThe sorptivity test was performed as per ASTM C1585-04.28

The test evaluated the increase in the mass of a 50 x 50 x 50 mm (1.97 x 1.97 x 1.97 in.) cubic specimen at given intervals of time (up to 360 minutes for initial sorptivity and up to 8 days for secondary sorptivity) when permitted to absorb water by capillary suction. Only the bottom surface of the specimen was in contact with water. The water depth was up to 4 mm (0.16 in.) to prevent water ingress from the sides; the perim-eter and top surface of the specimens were sealed with adhe-sive aluminum tape. This test was chosen because it measures the rate of ingress of water through unsaturated concrete and can therefore be considered a measure of water transport associated with capillary suction. Three specimens were used to determine the ingress of water for each ECC mixture.

Chloride-ion permeabilityThe chloride-ion permeability test, conducted in accor-

dance with ASTM C1202-97,29 measures the ease with which the charge passes through concrete, giving an indi-cation of the ECC’s resistance to chloride-ion permeability. Disc specimens 100 mm (4 in.) in diameter and 50 mm (2 in.) thick were cut from the midportion of 100 x 200 mm (4 x 8 in.) cylinder specimens and conditioned according to ASTM C1202.29 Specimens were then subjected to 60 V potential for 6 hours and the total charge that passed through the specimens was determined and used to evaluate the chlo-ride permeability of each ECC mixture. A minimum of three specimens were tested for each mixture.


MK were slightly higher than in the ECC mixtures produced with only FA. However, the midspan beam deflection values of the FA and MK mixtures (ECC-3 and ECC-4) were lower than those of the mixtures with only FA (ECC-1 and ECC-2). The midspan deflection values of the bending test demonstrated that the most important feature of ECC—high ductility with multiple-cracking behaviors—was maintained at an adequate level by replacing FA with 12.5% MK.

After the four-point bending test, the bending load was released and the specimens were taken out of the closed-loop controlled servo-hydraulic material test system. A crack closure occurred in the unloading position; the crack width in the loaded position was approximately 30% greater than in the unloaded position. All crack width measurements were conducted in the unloaded state. Crack widths were measured on the tension surface of the specimens using a portable crack microscope with 5 mm (0.00019 in.) magni-fication. Table 4 also shows the average crack widths and numbers on the span length of 102 mm (4.02 in.) at the center of the prism specimens’ tension surface. Each data point in Table 4 is an average of at least three or more prism specimens; more than 10 mm (0.00039 in.) crack widths were measured from each specimen. All four ECC mixtures showed crack widths of smaller than 75 mm (0.003 in.). Mixture ECC-2_FA/PC = 2.2 showed a very tight average crack with a width of 51 mm (0.002 in.). It was found that the number of cracks increased, whereas crack width decreased as FA content increased from 55 to 70%. ECC mixtures (ECC-3 and ECC-4) incorporating MK and FA led to a slightly wider crack width and a lesser number of cracks compared to FA-ECC mixtures.

Drying shrinkageThe results of drying shrinkage testing at 120 days after

the first 28 days of curing are provided in Fig. 2. The drying

RESULTS AND DISCUSSIONCompressive and flexural strengths and crack behaviors

The compressive strength variation of ECC mixtures is presented in Table 3. It shows that with increases in FA content and decreases in cement, compressive strength did not alter significantly from Mixtures ECC-1_FA/PC = 1.2 to ECC-2_FA/PC = 2.2. As seen in Table 3, the compres-sive strength of ECC mixtures incorporating FA and MK (Mixtures ECC-3_(FA + MK)/PC = 1.2 and ECC-4_(FA + MK)/PC = 2.2) was 20.3% and 12.8% higher than in the control ECC mixtures containing only FA (Mixtures ECC-1_FA/PC = 1.2 to ECC-2_FA/PC = 2.2), respectively. Inclusion of MK into the matrix improved the bond between the cement paste and silica sand particles and increased the density of the cement paste, which in turn significantly improved the compressive strength of the ECC mixtures.

Figure 1 shows the typical flexural-strength-midspan-beam deflection curves of the ECC mixtures. The bending capacity and flexural strength of these specimens are summa-rized in Table 4, which shows that the average ultimate flex-ural loads varied from 8.57 to 11.01 MPa (1.24 to 1.60 ksi) and the midspan beam deflection of the ECC beams at peak bending load varied from 4.30 to 7.17 mm (0.169 to 0.28 in.), depending on the content of FA or the FA/MK combination. Table 4 shows that increasing the FA/PC ratio from 1.2 to 2.2 (Mixtures ECC-1_FA/PC = 1.2 to ECC-2_FA/PC = 2.2) improved the bending deformation capacity by approxi-mately 33.5% while decreasing the flexural strength by approximately 22.3%. The improvement in bending defor-mation capacity with increased FA content can be attributed to the fact that greater amounts of FA tend to reduce the PVA fiber-matrix interface chemical bond and matrix toughness and increase the interface frictional bond in favor of attaining high bending capacity12,30 due to the change of matrix chem-ical composition and coating effect of inert particles on a fiber surface. Flexural strength test results also showed that the load-carrying capacities of ECC mixtures with FA and

Table 3—Compressive strength and chloride-ion permeability test results of ECCs

Mixture IDCompressive strength, MPa

Chloride-ion permeability

Coulombs Rating

ECC-1_FA/PC = 1.2 46.4 1072 Low

ECC-2_FA/PC = 2.2 46.8 1719 Low

ECC-3_(FA + MK)/PC = 1.2

55.5 627 Very low

ECC-4_(FA + MK)/PC = 2.2

52.8 1468 Low

Notes: 1 MPa = 1.6845 lb/yd3; 1 MPa = 0.145 ksi.

Fig. 1—Typical flexural-strength-midspan-deflection behavior of ECCs.

Table 4—Number of cracks, average crack widths, and bending test results of ECCs

Mixture ID

Bending test results After bending test

Midspan deflection at ultimate load, mm Flexural strength, MPa Number of cracks Residual crack width, mm

ECC-1_FA/PC = 1.2 5.37 11.01 33 65 ± 11

ECC-2_FA/PC = 2.2 7.17 8.57 41 51 ± 9

ECC-3_(FA + MK)/PC = 1.2 4.35 11.33 27 73 ± 12

ECC-4_(FA + MK)/PC = 2.2 4.30 8.75 32 68 ± 5

Notes: 1 MPa = 0.145 ksi; 1 mm = 0.0393 in.; 1 mm = 0.0000393 in.


shrinkage values of the ECC mixtures, as seen in the figure, ranged from 990 to 1450 me at 120 days. When the FA/PC ratio was increased from 1.2 to 2.2, drying shrinkage was reduced by approximately 14%. Yang et al.11 studied the effect of the FA/PC ratio on the drying shrinkage of high-volume FA-incorporated ECC and noticed that increasing the FA/PC ratio from 1.2 to 5.6 effectively decreased the drying shrinkage up to 50%. According to their conclu-sions, a possible mechanism behind the reduction of drying shrinkage in high-volume FA ECCs is the densification of the matrix, which may prevent internal moisture evapora-tion. Densification is typically attributed to the shape, pozzo-lanic property, and microfiller effect of FA. An alternative explanation would be that unhydrated FA particles serve as fine aggregates to restrain shrinkage deformation. The influence of MK incorporation on the drying shrinkage of ECC can also be seen in Fig. 2. The substitution of 10% (for Mixture ECC-3_(FA + MK)/PC = 1.2) and 12.5% (for Mixture ECC-4_(FA + MK)/PC = 2.2) MK with FA led to a reduction of 29.7% (according to Mixture ECC-1_FA/PC = 1.2) and 20.8% (according to Mixture ECC-2_FA/PC = 2.2) at the age of 120 days, respectively. The reduction in drying shrinkage with the incorporation of MK can be partly attributed to the lower amount of evaporable water, as the hydration and pozzolanic reaction used up a signifi-cant amount of the free water.7,31 With the inclusion of 10% MK for Mixture ECC-3_(FA + MK)/PC = 1.2 and 12.5% for ECC-4_(FA + MK)/PC = 2.2, the drying shrinkage of these two ECC mixtures became close to each other at 120 days. Mixtures ECC-3 and ECC-4 exhibited drying shrinkages of 1020 × 10–6 and 990 × 10–6, respectively, at the end of 120 days.

WA and WPFigure 3 presents the results of the WA and WP tests. An

increase of FA/PC from 1.2 to 2.2 remarkably increased both WA and WP. Mixture ECC-2_FA/PC = 2.2 had 8.63 and 14.89% WA and WP values, respectively, while those values were 6.36 and 10.34% for Mixture ECC-1_FA/PC = 1.2. It should be noted that the compressive and flexural strengths of ECC decreased (refer to Tables 3 and 4), while WP and WA increased. Similar findings for the mortar and concrete specimens have also been reported by other investigators. The most probable reason for the higher WA and WP with high volumes of FA is the slow pozzolanic reaction of FA due to an insufficient curing period. As explained in the “Specimen preparation and testing” section of this paper, the ECC specimens were cured in air after a 7-day sealed curing. With the incorporation of the blend of MK and FA (Mixtures ECC-3 and ECC-4), the WP and WA of the ECC mixtures improved due to an increased packing density. For example, the WP and WA values of the ECC mixture with 12.5% MK (ECC-4_(FA + MK)/PC = 2.2) decreased from 8.63% to 6.53% and from 14.89% to 12.84%, respec-tively. It is widely accepted that the principal reaction is facilitated by the dissolution of glassy/amorphous silica in pore water, which then reacts with CH to form CSH gel. The dissolution rate depends on the specific surface area, which is the main factor behind the strength, porosity, and pore diameter of various pozzolanic materials. Due to the relatively high specific surface area of MK (13,900 cm2/g [2154 in.2/g]), more of the silica enters the solution faster than FA (3870 cm2/g [600 in.2/g]), forming additional CSH gels on reaction and leading to an enhanced microstruc-

ture and a decreased value of the total porosity and WA of ECC.32 Khatip and Wild33 studied the pore size distribution of MK paste containing up to 15% MK and observed that the rate of pore refinement was very rapid up to 14 days of curing, after which the pore size changed slightly. This finding explains why the FA/MK mixtures had lower WA and porosity values than those containing only FA.

SorptivitySorptivity is a material property that characterizes the

tendency of a material to absorb and transmit water by capillary suction. Sorptivity testing measures the rate of capillary suction at a specified time and the sorptivity value indicates water mass uptake by concrete from the bottom surface.34 When testing the 50 mm (1.97 in.) cubic specimens, the cumulative WA per unit area up to 6 hours and 8 days was performed using linear regression analysis and the slope of equation was obtained to describe the initial and secondary sorptivity of the ECC mixtures, respectively. Figure 4 demonstrates the initial and secondary sorptivity coefficients. As seen in the figure, increasing the FA/PC ratio from 1.2 (Mixture ECC-1) to 2.2 (Mixture ECC-2) slightly increased the initial and secondary sorptivity coefficients of the ECC incorporating only FA. The initial and secondary sorptivity coefficients of Mixture ECC-1 (FA/PC = 1.2 and 55% FA content) were 0.0219 mm/sn0.5 and 0.0021 mm/sn0.5, respectively. However, these coefficients increased to 0.0331 mm/sn0.5 and 0.0024 mm/sn0.5 in Mixture ECC-2 (FA/PC = 2.2 and 70% FA content), respectively. A similar trend for mortar and ECC has also been observed by previous researchers.12,35 However, even at approximately a 70%

Fig. 2—Drying shrinkage variation of ECC mixtures.

Fig. 3—Water absorption and porosity test results of ECCs.


time of curing, most of the FA particles in the matrix expe-rienced no hydration and pozzolanic reactions. Because Mixture ECC-2 had more FA content (70%) than ECC-1 (55%), it was more negatively affected by the short period of curing. The benefits of using Class F FA to improve dura-bility properties, such as chloride-ion permeability resis-tance, are usually manifested at later ages with the contin-uous supply of moisture.12 Another possible reason could be that the fineness of the FA (3060 cm2/g [474 in.2/g]), as shown in Table 1, was significantly lower than the fineness of the cement used (3870 cm2/g [585 in.2/g]). As mentioned previously, increasing the fineness of cementitious materials positively affected the resistance of composites to chlo-ride-ion permeability. ECC mixtures (ECC-3 and ECC-4) produced with FA and MK had considerably lower chloride-ion permeability values than the ECC mixtures made only with FA. For instance, with the introduction of 10% MK, chloride-ion permeability decreased from 1072 coulombs (for Mixture ECC-1_FA/PC = 1.2) to 627 coulombs (for Mixture ECC-3_(FA + MK)/PC = 1.2), representing a reduction of approximately 42%. Mixtures ECC-2 (FA/PC = 2.2) and ECC-4 ((FA + MK)/PC = 2.2) demonstrated the same behavior. The binary use of FA and MK in Mixture ECC-4 (12.5% MK and 56% FA) decreased the chloride-ion permeability value of Mixture ECC-2 (70% FA) from 1719 to 1468 coulombs. Reduced capillary pores and reduced connectivity due to the rapid pozzolanic activity of MK, better particle packing, and higher Blaine fineness may be the reasons behind the better performance of ECCs with MK. Using FA and MK together can compensate for some of the shortcomings of ECC made exclusively with FA and create ECCs with increased durability.

CONCLUSIONSThe following conclusions can be drawn from this experi-

mental study:• An increased FA to PC ratio (FA/PC) did not signifi-

cantly alter the compressive strength of ECC mixtures produced with FA. However, the use of a binary blend of FA and MK in ECC production had a positive effect on the compressive strength; it increased from approxi-mately 12 to 20% with respect to ECC with only FA.

• Under the four-point bending test, all ECC mixtures exhibited multiple-cracking and strain-hardening behavior. Although the binary incorporation of FA and MK slightly decreased the midspan beam deflec-tion capacity of ECC specimens, it could still attain a capacity of up to 4.30 mm (0.169 in.)—significantly higher than that of normal concrete. Moreover, ECC mixtures with the binary use of FA and MK had some-what higher average crack width and flexural strength values than their ECC counterparts with only FA.

• As a result of the densification of the matrix and/or the unhydrated FA constraint effect, increasing the amount of FA reduced drying shrinkage by approximately 14%. The binary use of FA and MK in ECC produc-tion had a very positive effect on the drying shrinkage. The substitution of 10% (for Mixture ECC-3_(FA + MK)/PC = 1.2) and 12.5% (for Mixture ECC-4_(FA + MK)/PC = 2.2) MK with FA resulted in a reduction in the drying shrinkage as high as 30% and 21% for Mixtures ECC-1_FA/PC = 1.2 and ECC-2_FA/PC = 2.2 at 120 days, respectively.

replacement of cement with FA (Mixture ECC-2_FA/PC = 2.2), the initial sorptivity was still lower than the sorptivity coefficient of normal concrete. According to Neville,36 typical sorptivity is 0.09 mm/minute0.5 (0.00354 in./minute0.5) for normal concrete. Incorporating MK with FA in ECC produc-tion (Mixtures ECC-3 and ECC-4) positively affected the pore structure of the mixtures and significantly decreased both the initial and secondary sorptivity coefficients. Incorporating 10% MK with FA in Mixture ECC-3 ((FA + MK)/PC = 1.2) reduced the initial sorptivity coefficient from 0.0331 mm/sn0.5 (for Mixture ECC-1_FA/PC = 1.2) to 0.0224 mm/sn0.5 (for Mixture ECC-3_(FA + MK)/PC = 1.2). This trend was seen between Mixtures ECC-2 and ECC-4. These results show the value of using a binary FA/MK mixture rather than just FA on its own. The reduced sorptivity coefficient reflects a finer and impermeable pore structure that will, for example, inhibit ingress of aggressive agents into the pore structure.37

Chloride-ion permeabilityThe rapid chloride-ion permeability test results of the

ECC mixtures and their chloride-ion ratings according to ASTM C120229 are presented in Table 3. Rapid chloride-ion permeability testing is based on the electrical conductivity of ECC. The ECC sample is subjected to a potential differ-ence of 60 V and the total charge passing through it at the end of 6 hours is measured and expressed in coulombs. A reduction in this total charge value indicates better resistance to chloride-ion permeability and lower permeability.38,39 As seen in the table, increasing the FA/PC ratio from 1.2 (Mixture ECC-1 with 55% FA) to 2.2 (Mixture ECC-2 with 70% FA) reduced resistance to chloride-ion permeability. This result is surprisingly contrary to the findings of previous research performed on mortar and concrete. Normally, concrete with high volumes of pozzolans shows lower chloride-ion permeability due to a denser microstructure. The pozzolanic reaction may result in fewer capillary pores and less clog-ging of those pores, which reduces chloride-ion transport in concrete.40 The literature also mentions that the fineness of pozzolans has a great influence on chloride-ion perme-ability and, therefore, high fineness may have contributed to the lower chloride-ion permeability.41 As mentioned by Sahmaran and Li,12 however, the trend in ECC is completely different than in mortar and concrete. As seen in Table 3, increasing FA content from 55% (Mixture ECC-1_FA/PC = 1.2) to 70% (Mixture ECC-2_FA/PC = 2.2) increased chloride-ion permeability from 1072 to 1719 coulombs. The possible reason behind the higher chloride permeability with higher FA content is that the ECC specimens were cured in air after a 7-day sealed curing. Due to a relatively short

Fig. 4—Initial and secondary sorptivity test results of ECCs. (Note: 1 mm/minute0.5 = 0.0393 in./minute0.5.)


• Increasing the amount of FA in the ECC mixtures worsened their durability-related properties. Remark-able increases in WA, porosity, initial and secondary sorptivity, and chloride-ion permeability values were monitored. This can be attributed to the inadequate curing and relatively low FA fineness. With the use of a binary blend of FA and MK in ECC, however, all of the aforementioned durability-related properties improved significantly. This can be associated with reduced capil-lary pores and the reduction in pore connectivity due to the rapid pozzolanic reaction and higher Blaine fine-ness of MK, as well as better particle packing density of the matrix.

ACKNOWLEDGMENTSThe authors gratefully acknowledge the financial assistance of The

Council of Higher Education of Turkey (YOK), the Natural Sciences and Engineering Research Council (NSERC) of Canada, and the Canada Research Chair Program.

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24. Sahmaran, M.; Keskin, S. B.; Ozerkan, G.; and Yaman, I. O., “Self-Healing of Mechanically Loaded Self-Consolidating Concretes with High Volumes of Fly Ash,” Cement and Concrete Composites, V. 30, 2008, pp. 872-879.

25. ASTM C39-94, “Standard Test Method for Compressive Strength of Cylindrical Concrete Specimens,” ASTM International, West Conshohocken, PA, 1994.

26. ASTM C157/C157M-04, “Standard Test Method for Length Change of Hardened Hydraulic-Cement, Mortar, and Concrete,” ASTM Interna-tional, West Conshohocken, PA, 2004, 7 pp.

27. ASTM C642-06, “Standard Test Method for Density, Absorption, and Voids in Hardened Concrete,” ASTM International, West Conshohocken, PA, 2006, 3 pp.

28. ASTM C1585-04, “Standard Test Method for Measurement of Rate of Absorption of Water by Hydraulic-Cement Concretes,” ASTM Interna-tional, West Conshohocken, PA, 2004, 6 pp.

29. ASTM C1202-97, “Standard Test Method for Electrical Indication of Concrete’s Ability to Resist Chloride Ion Penetration,” ASTM International, West Conshohocken, PA, 1997, 6 pp.

30. Wang, S., and Li, V. C., “Engineered Cementitious Composites with High-Volume Fly Ash,” ACI Materials Journal, V. 104, No. 3, May-June 2007, pp. 233-241.

31. Brooks, J. J., and Megat Johari, M. A., “Effect of Metakaolin on Creep and Shrinkage of Concrete,” Cement and Concrete Composites, V. 23, 2001, pp. 495-502.

32. Sabir, B. B.; Wild, S.; and Bai, J., “Metakaolin and Calcined Clay as Pozzolans for Concrete: A Review,” Cement and Concrete Composites, V. 23, 2001, pp. 441-454.

33. Khatip, J. M., and Wild, S., “Pore Size Distribution of Metakaolin Paste,” Cement and Concrete Research, V. 26, No. 10, 1996, pp. 1545-1553.

34. Siddique, R., and Kaur, A., “Effect of Metakaolin on the Near Surface Characteristics of Concrete,” Materials and Structures, V. 44, 2011, pp. 77-88.

35. Chindaprasirt, P.; Jaturapitakkul, C.; and Sinsiri, T., “Effect of Fly Ash Fineness on Compressive Strength and Pore Size of Blended Cement Paste,” Cement and Concrete Composites, V. 27, 2005, pp. 425-428.

36. Neville, A. M., Properties of Concrete, fourth edition, Longman Group Limited, New York, 1995, 844 pp.

37. Bai, J.; Wild, S.; and Sabir, B. B., “Sorptivity and Strength of Air-Cured and Water-Cured PC–PFA–MK Concrete and the Influence of Binder Composition on Carbonation Depth,” Cement and Concrete Research, V. 32, 2002, pp. 1813-1821.

38. Sengul, O., and Tasdemir, M. A., “Compressive Strength and Rapid Chloride Permeability of Concretes with Ground Fly Ash and Slag,” Journal of Materials in Civil Engineering, ASCE, V. 21, No. 9, 2009, pp. 494-501.

39. Sengul, O.; Tasdemir, C.; and Tasdemir, M. A., “Influence of Aggregate Type on Mechanical Behavior of Normal- and High-Strength Concretes,” ACI Materials Journal, V. 99, No. 6, Nov.-Dec. 2002, pp. 528-533.

40. Li, S., and Roy, D. M., “Investigation of Relations between Porosity, Pore Structure and Chloride Diffusion of Fly Ash and Blended Cements,” Cement and Concrete Research, V. 16, No. 5, 1986, pp. 749-759.

41. Zhang, M. H.; Bilodeau, A.; Malhotra, V. M.; Kim, K. S.; and Kim, J. C., “Concrete Incorporating Supplementary Cementing Materials: Effect on Compressive Strength and Resistance to Chloride Ion Penetration,” ACI Materials Journal, V. 96, No. 2, Mar.-Apr. 1999, pp. 181-189.


Title no. 109-M56




Fatigue Analysis of Plain and Fiber-Reinforced Self-Consolidating Concreteby S. Goel, S. P. Singh, and P. Singh

SCC possesses good fluidity and deformability, making it more suitable for the addition of fibers as compared to NVC and allows for much easier construction, resulting in a more reliable quality in concrete placement and a more homogeneous material structure.14 SCC reinforced with steel fibers enhances its applications because the mechanical performance of concrete is improved. Self-consolidating fiber-reinforced concrete (SCFRC) is more ductile and tougher than conventional SCC and has demonstrated higher residual strengths.15 The workability of SCFRC is directly influenced by the type and content of fibers used, as well as the SCC matrix. A higher aspect ratio and volume concen-tration of fibers improve the performance of SCFRC in the hardened state but also affects its workability. Thus, studies were conducted to obtain optimum fiber-reinforced concrete (FRC) mixtures with required self-consolidating proper-ties.16-18 Dhonde et al.17 revealed that SCFRC could be made with satisfactory filling and passing ability using short fibers (L ≤ 30 mm [1.2 in.] long), as these did not influence its slump flow or stability. Researchers investigated whether SCFRC shows either similar or improved performance in terms of compressive strength, flexural strength, splitting tensile strength, elastic modulus, creep and shrinkage, and shear and pullout behavior compared to SCC and normally vibrated fiber-reinforced concrete (NVFRC) under statically applied loads.11,14,15,17,19-21 The microstructure around the matrix, the distribution, and the orientation of the fibers are different in SCC than in conventional concrete. Entrapped air and neighboring fibers affect the performance of a fiber in NVFRC more than in SCFRC. The steel fibers, due to the lack of any mechanical vibrations in SCFRC, are more favorably aligned into the direction of the flow, thereby improving its bending characteristics. In SCC, the fibers are fully embedded in the matrix, thereby imparting better bond or pullout strength.13,21 Thus, it is expected that SCC and SCFRC, as in the case of their mechanical properties, such as compressive and flexural strength under statically applied loads, may exhibit better fatigue characteristics.

The global thrust on construction of bridges and highway pavements for infrastructure development has fascinated many researchers,3,4,7,8 leading to investigations of the fatigue behavior of concrete. The bridges and pavements were expected to resist millions of cycles of repeated axle loads during their intended life. Considering fatigue strength an important parameter in the design of these structures,

This paper investigates the flexural fatigue performance of self-consolidating concrete (SCC) and self-consolidating fiber-reinforced concrete (SCFRC) containing round corrugated steel fibers with a size of 1 x 30 mm (0.04 x 1.18 in.) in different 0.5, 1.0, and 1.5% volume fractions. Approximately 250 flexural fatigue tests and 195 complementary static flexural tests were executed on beam specimens with a size of 100 x 100 x 500 mm (3.94 x 3.94 x 19.7 in.) under four-point flexural loading. The fatigue-life data show that the probabilistic distribution of fatigue life of SCC/SCFRC at a given stress level can approximately be modeled by the two-parameter Weibull distribution. Three different methods were used to obtain the Weibull parameters. A single-log fatigue equation was used to analyze the flexural fatigue performance of SCC/SCFRC with a 10% probability of failure. The results show significantly improved fatigue performance of SCFRC with enhanced sensitivity of fatigue lives to the change of applied stress. Theoretic fatigue lives for SCC/SCFRC were estimated that exhibit an increase to a different extent.

Keywords: fatigue life; self-consolidating fiber-reinforced concrete; stress level; Weibull distribution.

INTRODUCTIONSelf-consolidating concrete (SCC) is an innovative concrete

that does not require vibration for placing and compaction. It is able to flow under its own weight, completely filling formwork, and encapsulate the reinforcement, achieving full compaction, even in the presence of congested reinforce-ment.1 The hardened SCC is dense and homogeneous and has improved engineering properties and durability compared to normally vibrated concrete (NVC). The improved construc-tion practice and performance, combined with the health and safety benefits, make SCC a very attractive solution for both precast concrete and civil engineering construction.2

Due to its substantial engineering applications and commercial benefits, SCC has generated tremendous interest among researchers, engineers, and concrete technolo-gists.3,4 Numerous research studies have shown that it is prac-tical to make a flowable yet stable SCC tailored for any appli-cation.5,6 A number of investigations related to the rheological, mechanical, and structural behavior of SCC under statically applied loads have been reported in literature that substantiate the better performance of SCC compared to NVC.7-10

The importance of the homogeneity of the material is evident for any application because it will affect the material properties.11 SCC contains large proportions of finer parti-cles and does not need mechanical vibrators for compaction, which results in a denser and more homogenous concrete compared to NVC. The denser structure of SCC dimin-ishes the presence of air voids so better bonding between the concrete and reinforcing materials is achieved; this could be beneficial and lead to better results in terms of the mechanical behavior of the constituents compared with that of conventional concrete.12,13


behavior of SCC/SCFRC. To this end, an experimental investigation was set up to establish the probability distri-butions for fatigue/fatigue-life data of SCC and SCFRC at different stress levels. The two-parameter Weibull distribu-tion was examined in this regard and distribution parameters were obtained and compared with that of NVC and NVFRC. To examine the fatigue performance, the Weibull distribu-tion was used to incorporate the probability of fatigue failure into the fatigue-life data and the theoretic fatigue lives for SCC and SCFRC were obtained and compared with those of NVC and NVFRC.

EXPERIMENTAL INVESTIGATIONMaterials and mixture proportions

The concrete mixtures were prepared with Grade 43 ordinary portland cement conforming to Indian Standard (IS) 8112 and fly ash (Class F). The mixtures were prepared using well-graded crushed stone coarse aggregate with a nominal size of 12.5 mm (0.49 in.) and locally available coarse sand with a fineness modulus of 2.85. A polycarboxylic-ether-based high-range water-reducing admixture (HRWRA) and a polycarboxylate-polymer-based viscosity-modifying agent (VMA) were used to achieve the flowable yet cohesive SCC and SCFRC mixtures. Corrugated steel fibers were 30 mm (1.18 in.) in length and 1 mm (0.04 in.) in diameter in all the SCFRC mixtures. Table 1 shows the proportions of all four mixtures of SCC and SCFRC used in this investiga-tion. The mixture with no steel fibers—that is, the SCC mixture—was taken as the control mixture. Three different SCFRC mixtures contained steel fibers in volume fractions of 0.5, 1.0, and 1.5%. The dosage of HRWRA and VMA was adjusted to obtain the required workability for all the SCC and SCFRC mixtures. All the mixtures were mixed in a 100 L (0.1 m3) drum mixer in the laboratory. First, the fine and coarse aggregates were fed into the mixer and mixed for approximately 1 minute. The cement and fly ash were added to the aggregates and the ingredients were mixed in a dry condition for approximately 30 seconds. Subsequently, two-thirds of the water was added to the dry mixture and mixing was allowed for the next 60 seconds. HRWRA premixed with the remaining one-third of the water was added to the wet mixture and mixing continued for another 150 seconds. In the case of the SCFRC mixtures, the steel fibers were added to the wet mixture by uniformly sprinkling them into the drum, and then the remaining one-third of the water premixed with HRWRA and VMA was added. Mixing was allowed for another 60 seconds for the SCFRC mixtures. The SCC mixture did not show any sign of bleeding but the SCFRC mixtures were unstable and bleeding was observed during the filling of the molds; thus, a polycarboxylate-polymer-based VMA was used to improve the stability of the SCFRC mixtures.

S. Goel is an Assistant Professor in the Department of Civil Engineering at DAV Insti-tute of Engineering and Technology, Jalandhar, India. His research interests include self-consolidating concrete and recycling of materials in concrete.

ACI member S. P. Singh is a Professor in the Department of Civil Engineering at Dr. B. R. Ambedkar National Institute of Technology, Jalandhar, India. He received his PhD from the University of Roorkee, Roorkee, Uttarakhand, India, in 1999. His research interests include the fatigue behavior of fibrous concrete composites and recycling of materials in concrete.

P. Singh is a Professor in the Department of Civil Engineering at Dr. B. R. Ambedkar National Institute of Technology. He received his PhD from Panjab University, Chan-digarh, India, in 2002. His research interests include the behavior of laminated plates and fiber-reinforced concrete.

the majority of research in the literature on the fatigue of NVC and NVFRC has been focused on its behavior in flexure.22-25 The research investigations were carried out to suggest relationships between stress level S, which is the ratio of maximum fatigue stress fmax to the modulus of rupture fr and the number of load cycles N that causes failure. One of the extensively used fatigue equations is a single-log equation, as shown in the following22,25-27

10log ( )max

r

fS a b N

f= = − (1)

where a and b are experimental coefficients. The fatigue test data of NVC and NVFRC are random in nature and show considerable scatter; even under carefully controlled test procedures; thus, it becomes essential to introduce probabilistic concepts to ensure adequate fatigue resis-tance. Oh,27,28 Singh and Kaushik,25,29 and Mohammadi and Kaushik30 conducted experimental and theoretical studies to investigate the fatigue-life distributions of NVC and NVFRC at different stress levels. From the test data, it was observed that the statistical distribution of the fatigue life of NVC and NVFRC can be approximately described by the two-param-eter Weibull distribution.28-30 A number of investigations have been made to study the fatigue behavior of NVC and NVFRC; the fatigue characteristics of SCC/SCFRC have yet to be investigated despite the fact that SCC/SCFRC has been widely accepted for the construction of bridge deck, bridge piers, and pier caps4 and possible future applications in highway and airfield concrete pavements, wherein fatigue is the predominant mode of loading.

RESEARCH SIGNIFICANCEThe literature available on SCC and SCFRC reported

studies on their mechanical properties under statically applied loads, but to the best of the authors’ knowledge, there is practically no information available on the fatigue

Table 1—Proportions for SCC and SCFRC mixtures

MixtureCement,

kg/m3 (lb/yd3)Fly ash,

kg/m3 (lb/yd3)Fine aggregates, kg/m3 (lb/yd3)

Coarse aggregates, kg/m3 (lb/yd3)

Fiber-volume fraction, Vf

HRWRA, by weight of cement

VMA, by weight of cement

SCC 410 (691) 205 (346) 846 (1427) 602 (1015) NA 1.7% NA

SCFRC0.5 410 (691) 205 (346) 846 (1427) 602 (1015) 0.5% 1.9% 0.25%

SCFRC1.0 410 (691) 205 (346) 846 (1427) 602 (1015) 1.0% 2.2% 0.35%

SCFRC1.5 410 (691) 205 (346) 846 (1427) 602 (1015) 1.5% 2.5% 0.50%

Note: NA is not available.


Workability and casting of specimensAll the workability tests were conducted after approx-

imately 1 minute of final mixing; namely, slump flow, V-funnel, J-ring, and L-box tests were carried out for SCC and SCFRC mixtures per the guidelines of EFNARC.2 The results of the workability tests conducted randomly on five batches each of SCC and SCFRC mixtures are presented in Table 2.

Standard beam specimens with a size of 100 x 100 x 500 mm (3.94 x 3.94 x 19.7 in.) for static flexural and flexural fatigue tests and cube specimens with a size of 150 x 150 x 150 mm (5.9 x 5.9 x 5.9 in.) for compressive strength tests were cast in different batches. Each batch contained seven beam and three cube specimens. The mixture was poured into the specimen molds from a height of approximately 450 mm (17.7 in.) in a single layer. Molds were filled without any

use of vibrator. The specimens were demolded 36 hours after casting and moist-cured under laboratory conditions. For ascertaining the quality of each batch of SCC and SCFRC, compressive strength tests were conducted on cube specimens after 28 days of curing. The beam specimens were cured for 75 days and thereafter stored under laboratory conditions for approximately 2 months to minimize the effect of strength gain during the course of fatigue testing, which in itself is extended in nature. To further eliminate the effect of strength gain, if any, the testing was done batch-wise, wherein the static flexural strength tests on a particular batch were conducted just prior to the flexural fatigue testing of the same. The results of the compressive strength tests on SCC and SCFRC speci-mens are reported in Table 3. The average 28-day compressive strength for all batches of SCC, SCFRC0.5, SCFRC1.0, and

Table 2—Workability tests on fresh SCC and SCFRC mixtures

Test Parameter SCC SCFRC0.5 SCFRC1.0 SCFRC1.5 EFNARC guidelines

Slump flow

T500*, seconds 2.8 ± 0.5 3.0 ± 0.5 3.5 ± 0.5 4.0 ± 0.5 2 to 5

Slump flow spread*, mm (in.)

750 ± 20 (29.5 ± 0.8)

710 ± 30 (27.9 ± 1.2)

700 ± 30 (27.6 ± 1.2)

700 ± 20 (27.6 ± 0.8)

650 to 800 (25.6 to 33.5)

J-ring

T500J, seconds 3.0 ± 0.5 4.0 ± 0.5 4.0 ± 0.5 5.0 ± 0.5 3 to 6

Flow spread*, mm (in.)

720 ± 25 (28.3 ± 1)

710 ± 25(27.9 ± 1)

700 ± 20 (27.6 ± 0.8)

680 ± 20 (26.8 ± 0.8)

600 to 750 (23.6 to 29.5)

Blocking step* Bj, mm (in.)

6.0 ± 0.5(0.24 ± 0.02)

7.0 ± 0.5(0.3 ± 0.02)

8.0 ± 0.5(0.33 ± 0.02)

9.5 ± 0.4(0.38 ± 0.16)

0 to 10 (0.4)

V-funnelV-funnel time*,

seconds7.0 ± 0.5 7.7 ± 0.3 8.5 ± 0.5 9.5 ± 0.5 6 to 12

L-boxL-box passing

ability* 0.91 0.90 0.83 0.81 0.8 to 1.0

*For random five batches. Note: 1 in. = 25.4 mm.

Table 3—Compressive strength test results for SCC and SCFRC mixtures

Batch No.

28-day average* compressive strength, MPa (psi)

SCC SCFRC0.5 SCFRC1.0 SCFRC1.5

1 35.9 (5210) 36.0 (5225) 41.1 (5965) 42.0 (6096)

2 35.2 (5109) 39.9 (5791) 39.1 (5675) 39.8 (5776)

3 35.0 (5080) 36.2 (5254) 39.6 (5747) 44.0 (6386)

4 34.9 (5065) 36.9 (5355) 41.6 (6038) 39.9 (5791)

5 36.6 (5312) 39.6 (5747) 41.2 (5980) 42.9 (6226)

6 35.7 (5181) 37.7 (5472) 39.3 (5704) 40.8 (5922)

7 36.7 (5326) 38.5 (5588) 39.8 (5776) 42.7 (6197)

8 35.2 (5109) 36.1 (5239) 41.6 (6038) 43.6 (6328)

9 36.8 (5341) 39.4 (5718) 39.8 (5776) 41.8 (6067)

10 36.5 (5298) 36.6 (5312) 42.2 (6125) 43.6 (6328)

11 35.8 (5196) 39.9 (5791) 38.8 (5631) 42.9 (6226)

12 36.5 (5297) 36.6 (5312) 40.8 (5922) 43.8 (6357)

13 36.3 (5268) 38.2 (5544) 39.0 (5660) 42.5 (6168)

14 35.5 (5152) 38.8 (5631) 38.9 (5646) 41.9 (6081)

15 36.2 (5254) 40.1 (5820) 41.0 (5951) 43.8 (6357)

16 35.4 (5139) 38.3 (5559) 40.7 (5907) 41.7 (6052)

Average 35.9 (5210) 38.1 (5530) 40.3 (5849) 42.4 (6154)*Average of three specimens. Note: 1000 psi = 6.89 MPa.


the fatigue stress ratio R (R = fmin/fmax), kept constant at 0.10 throughout the investigation, as has been done in previous studies.29-31 Constant-amplitude sinusoidal loads were applied at a frequency of 10 Hz to complete the test in a reasonable period of time. Because fatigue testing is a time-consuming and expensive process and a large number of specimens were proposed to be tested in this investiga-tion, an upper limit of 2 million cycles of fatigue loading was adopted. The test was terminated when the failure of the specimen occurred or this upper limit was reached, which-ever was earlier. For each SCC and SCFRC mixture, the numbers of cycles to failure for the specimen under different load conditions were noted as fatigue life N.

ANALYSIS OF FATIGUE-LIFE DATAThe fatigue test data obtained for the SCC and SCFRC

mixtures in this study shows considerable variability, even at a given stress level. Thus, before initiating the analysis process, some data points may deserve consideration for rejection as outliers. Chauvenet’s criteria32 was applied to the fatigue-life data at different stress levels tested in this investigation and points meeting this criterion for outliers were identified and excluded from further analysis. Batson et al.,33 Singh and Kaushik,29 and Mohammadi and Kaushik30 used the same criterion in their work on the flexural fatigue of plain NVC and NVFRC.

Fatigue-life distributions of SCC and SCFRCThe fatigue test data of concrete is known to exhibit great

variability, which becomes enhanced in the case of FRC and thereby necessitates introducing the probability concepts in the design to secure the adequate fatigue resistance of concrete structures such as bridges, highway pavements, and

SCFRC1.5 was 35.90, 38.10, 40.30, and 42.40 MPa (5210, 5530, 5849, and 6154 psi), respectively.

Fatigue test programThe flexural fatigue testing of SCC and SCFRC was the

primary objective of this investigation. The maximum and minimum load limits are required to be defined to initiate a fatigue test. These load limits were obtained for each batch of specimens by testing three beam specimens from a particular batch in static flexure. The beams were simply supported over a span of 450 mm (17.7 in.) and loaded at third points, thus leading to a four-point bending test. The average static flexural strength fr for each batch of SCC and SCFRC was obtained just before the fatigue tests, the results of which are presented in Table 4. The static flexural tests were carried out with a 100 kN (22.2 kip) servo-controlled actuator run in the displacement control mode at a loading rate of 0.5 mm/minute (0.02 in./minute). The static flex-ural strength taken as an average of all the batches of SCC, SCFRC0.5, SCFRC1.0, and SCFRC1.5 was 4.85, 6.05, 7.20, and 9 MPa (704, 878, 1045, and 1308 psi), respectively. A considerable increase in the peak loads over the first crack loads was observed for SCFRC specimens, particularly for mixtures containing 1.0 and 1.5% fiber-volume fractions. The increment in peak load may be attributed to the contri-bution of fibers after the cracking of the matrix.

After the static flexural strength of a particular batch of SCC or SCFRC was established, the remaining beam speci-mens were tested in flexural fatigue. The loading conditions were kept the same for both static flexural and flexural fatigue tests. The flexural fatigue tests were conducted at stress levels S (S = fmax/fr, fmax is maximum fatigue stress, and fr is static flexural strength), ranging from 0.90 to 0.65 with

Table 4—Static flexural strength test results for SCC and SCFRC mixtures

Batch No.

Static flexural strength*, MPa (psi)


1 4.93 (715) 6.23 (904) 6.48 (940) 9.23 (1340)

2 4.96 (720) 6.35 (922) 7.60 (1103) 9.18 (1332)

3 4.60 (668) 6.11 (887) 7.60 (1103) 9.41 (1366)

4 5.02 (729) 5.84 (848) 6.93 (1006) 9.06 (1315)

5 4.87† (707) 6.02 (874) 7.77 (1128) 8.78 (1274)

6 4.42† (642) 5.89 (855) 7.94† (1152) 8.89† (1290)

7 4.67 (678) 6.26 (909) 7.15 (1038) 8.45 (1226)

8 4.69† (681) 6.16 (894) 6.54 (949) 9.28 (1347)

9 5.43 (788) 5.82 (845) 7.44† (1080) 9.32 (1353)

10 4.82 (700) 5.96† (865) 6.98 (1013) 8.31 (1206)

11 4.42 (642) 6.36 (923) 7.91 (1148) 9.44 (1370)

12 4.64 (673) 5.74 (833) 7.86 (1141) 8.73† (1267)

13 5.38 (781) 5.66 (821) 6.72 (975) 8.60 (1248)

14 4.72 (685) 5.92 (859) 6.85 (994) 8.69 (1261)

15 5.10 (740) 6.28 (911) 6.69 (971) 9.40 (1364)

16 4.89 (710) 6.21 (901) 6.76 (981) 9.34 (1355)

Average 4.85 (704) 6.05 (878) 7.20 (1045) 9.00 (1308)*Results correspond to different ages at testing. †Average of two specimens. Notes: Without mark is average of three specimens; 1000 psi = 6.89 MPa.


offshore structures. In this study, it is proposed to use the two-parameter Weibull distribution to describe the proba-bility distributions of fatigue life of SCC and SCFRC. Unlike lognormal distribution suggested by ASTM 91-A34 that shows decreasing hazard function, the Weibull distribution gives increasing hazard function with an increase in life or time, which demonstrates the actual behavior of engineering materials subjected to fatigue load.32 Moreover, the Weibull distribution is based on more convincing arguments, is rela-tively easy to use, has well-developed statistics,35 and has been used in previous studies for the statistical description of fatigue data of NVC27,28 and NVFRC.29,30 In this study, the graphical method was employed to show that the distri-bution of fatigue life of SCC and SCFRC at a given stress level S follows the two-parameter Weibull distribution. To estimate the parameters of the Weibull distribution, different methods—that is, the graphical method, the method of moments, and the method of maximum likelihood esti-mate—were suggested.36 The calculation of parameters by different methods instills confidence in the results and hence has been adopted herein.

Graphical method and distribution parametersThe survivorship function LN(n) of the two-parameter

Weibull distribution may be written as follows27-30

expNnLu

α = − (2)

where n is the specific value of random variable N; a is the shape parameter at stress level S; and u is the scale parameter at stress level S.

Taking the logarithm twice on both sides of Eq. (2)

( )1ln ln ln ln( )N

n uL

= α − α

(3)

Equation (3) represents a linear relationship between ln[ln(1/LN)] and ln(N). To obtain a graph from Eq. (3), the fatigue-life data corresponding to a particular stress level S are first arranged in ascending order of cycles to failure, and the empirical survivorship function LN for each fatigue-life data is obtained from the following expression27-30

11N

iLk

= −+

(4)

where i denotes the failure order number; and k represents the number of fatigue data points in a data sample under consid-eration at a given stress level S. A graph is plotted between ln[ln(1/LN)] and ln(N), and if the test data, at a particular stress level, follow approximately a linear trend, then this indicates that the two-parameter Weibull distribution is a reasonable assumption for the statistical description of fatigue-life data at that stress level. The shape parameter a and the scale parameter u can be obtained either from the graph or directly from the regression analysis. Figure 1 shows the plot of the fatigue-life data of SCC at stress level S = 0.85, 0.80, 0.75, 0.70, and 0.65. The approximate straight-line plot indicates

that the two-parameter Weibull distribution is a reasonable assumption for the distribution of fatigue life of SCC at all tested stress levels. The corresponding values of corre-lation coefficient CC are 0.9768, 0.9919, 0.9912, 0.9926, and 0.9911 for fatigue-life data of SCC at stress levels of 0.85, 0.80, 0.75, 0.70, and 0.65, respectively. The values of the correlation coefficient for all stress levels were greater than 0.97, which further substantiated the validity of the two-parameter Weibull distribution for the fatigue life of SCC. The parameters a and u for all stress levels for the fatigue-life data of SCC were estimated directly from the regression analysis.

Similarly, the fatigue-life data of SCFRC with 0.5, 1.0, and 1.5% volume fractions of fibers at different stress levels were analyzed by the graphical method and were shown to follow the two-parameter Weibull distribution with the statistical correlation coefficient exceeding 0.90. Figures 2 to 4 present the plots of graphical analysis of SCFRC with 0.5, 1.0, and 1.5% volume fractions of fibers at different stress levels. The corresponding correlation coefficients are also listed in these figures. The parameters of SCFRC containing different fiber-volume fractions were estimated by this method.

Parameters from method of momentsThe estimation of parameters of the Weibull distribution

by the method of moments requires sample moments, such as sample mean and sample variance. The following rela-tionships can be used to obtain the parameters a and u for fatigue-life data of SCC and SCFRC28,29,36

a = (CV) –1.08 (5)

and

1 1u µ

= Γ + α

(6)

Fig. 1—Graphical analysis of fatigue-life data for SCC at different stress levels S.


q = ua (8)

The maximum likelihood equations can be modified as follows

( )( )

*

*

1*

1

1

ln 1 1 ln

k

i i ki

iki

ii

n nn

kn

α

=

α =

=

∑− = ∑

α∑ (9)

**

1

1 k

ii

nk

α

=θ = ∑ (10)

where a* and q* are the maximum likelihood estimates of a and q, respectively. The value of the shape parameter is first obtained by solving Eq. (9) by an iterative procedure. An estimate of the value of a obtained by any of the two preceding methods—that is, the graphical method and method of moments—can be used as a first trial. Once the shape parameter is known, the value of u can be obtained from Eq. (8).

The parameters of the Weibull distribution were also obtained by the method of maximum likelihood for fatigue-life data of SCC and SCFRC with 0.5, 1.0, and 1.5% of volume fraction of steel fibers corresponding to different stress levels S ranging from 0.90 to 0.65. The average values of the parameters for SCC and SCFRC obtained from different methods are listed in Table 5.

To analyze the beneficial effects of self-consolidation on the fatigue life of concrete, the results of this investigation for the fatigue life of SCC and SCFRC were compared with some previous studies on the fatigue of NVC and NVFRC. The work of Oh,27,28 Singh and Kaushik,25,29 and Mohammadi and Kaushik30 on NVC and NVFRC was selected for compar-ison with SCC and SCFRC, as the aggregate type and size; specimen size; static flexural strength of the concretes; and the shape, size, and volume fractions of the steel fibers used in these studies are comparable to those used in this study.

where m is the sample mean of the fatigue-life data at a given stress level; CV (= s/m, s is standard deviation of sample) is the coefficient of variation of the data; and G() is the gamma function. The parameters of the Weibull distribution for the fatigue-life data of SCC and SCFRC with 0.5, 1.0, and 1.5% of volume fraction of steel fibers were estimated using Eq. (5) and (6) for stress level S ranging from 0.9 to 0.65.

Parameters from method of maximum likelihood estimate

The distribution parameters of the Weibull distribution can also be obtained using the method of the maximum likelihood estimate. The probability density function of the Weibull distribution may be written as follows28-30

( ) 1 expNnf n n

αα− α

= − θ θ (7)

where

Fig. 2—Graphical analysis of fatigue-life data for SCFRC with 0.5% steel fibers at different stress levels S.




Because the fatigue-life data of NVC and NVFRC specimens of comparable size and static flexural strength were available, it was thought prudent to concentrate all efforts and resources on SCC and SCFRC and avoid repeating available results. This was also demanded by the economy of the investiga-tion because a large number of specimens were required to be tested for improving the reliability of widely scattering fatigue test results. The shape parameters obtained from this investigation for SCC and SCFRC, together with the values of shape parameters for NVC and NVFRC taken from the literature, are plotted in Fig. 5 through 7.

Figures 5 through 7 show that the shape parameter for the fatigue-life data of SCC and SCFRC decreases with a decrease in the stress level, thus indicating higher vari-ability in the fatigue-life distribution of SCC and SCFRC at lower stress levels. Similar results have been reported by previous investigators for the fatigue life of NVC28,30 and NVFRC.29,30 The relative magnitude of the shape parameter is an indicator of variability in the distribution of fatigue life such that a relatively higher value of the shape param-eter represents a lower variability in the distribution of fatigue life and vice versa. The plots in Fig. 5 through 7 and results compiled in Table 5 show that across all the consid-ered values of the shape parameter, there was a lower vari-

Table 5—Average values of Weibull parameters for fatigue life of SCC

Stress level, S


a u a u a u a u

0.90 — — 3.4676 2122 2.8216 6277 2.3446 1959

0.85 4.3471 2106 2.6992 10,378 1.9790 55,992 1.8575 22,852

0.80 3.7858 15,825 2.0985 41,073 1.7666 250,772 1.6634 123,637

0.75 3.3747 66,618 1.8556 246,028 1.5993 1,044,763 1.4917 505,512

0.70 3.1740 192,119 1.7491 961,500 — — 1.3830* 105,597*

0.65 2.9867 1,415,250 — — — — — —*Average of method of moment and method of maximum likelihood.

Fig. 5—Comparison of shape parameters for fatigue life of SCC, NVC, SCFRC, and NVFRC with Vf = 0.5% of steel fibers.

Fig. 6—Comparison of shape parameters for fatigue life of SCFRC and NVFRC with Vf = 1.0% of steel fibers.

Fig. 7—Comparison of shape parameters for fatigue life of SCFRC and NVFRC with Vf = 1.5% of steel fibers.


ability in the distribution of fatigue life of SCC and SCFRC compared to NVC and NVFRC, respectively. For example, at stress level S = 0.85, the average value of the shape param-eter for the fatigue life of SCC, obtained by the different methods of analysis in this investigation, is 4.3471 compared with 3.8920 and 3.5457 reported by Oh28 and Mohammadi and Kaushik,30 respectively, for NVC, as plotted in Fig. 5. This trend is established at all other stress levels tested in this study. In particular, the shape parameter of SCC is higher than that of NVC by 17%, 36%, 34%, 41%, and 36% at stress level S = 0.85, 0.80, 0.75, 0.70, and 0.65, respec-tively. A maximum decrease of approximately 15% in the coefficient of variation in the fatigue-life data of SCC was observed as compared to NVC.

Similar trends have been observed for fatigue life of SCFRC with 0.5, 1.0, and 1.5% volume fractions of steel fibers. The shape parameters calculated by all three aforemen-tioned methods for SCFRC were found to be greater than the shape parameters for NVFRC with the same volume fractions of steel fibers at different stress levels tested in the investiga-tion. For example, the average value of the shape parameter for the fatigue life of SCFRC with 0.5% volume fractions of steel fibers at stress level S = 0.80 was 2.0985, as compared with 1.5448 reported by Singh and Kaushik25 for NVFRC with 0.5% steel fibers. The average value of the shape param-eters for SCFRC with 1.0 and 1.5% volume fraction of steel fibers is 1.6904 and 1.6634 at stress level S = 0.80, compared to 1.2385 and 1.4376 reported by Singh and Kaushik29 and Mohammadi and Kaushik,30 respectively, for SFRC. The maximum increase of 37%, 42%, and 55% in the shape param-eter for SCFRC was observed compared to NVFRC with 0.5%, 1.0%, and 1.5% of steel fibers, respectively, reported by Singh and Kaushik29; at the same time, a maximum decrease of 21%, 25%, and 28% in the coefficient of variation for the fatigue-life data of SCFRC was observed with 0.5%, 1.0%, and 1.5% volume fraction of steel fibers, respectively. It can also be observed from Fig. 5 to 7 that the shape parameter decreases as the fiber-volume fraction increases, resulting in higher variability in the fatigue life of SCFRC at higher fiber-volume fractions, even at the same stress level. In general, for all mixtures of SCFRC with different volume fractions of steel

fibers, the values of the shape parameter are lower as compared to those of SCC. This indicates higher variability in the distri-bution of fatigue life of SCFRC, as compared to SCC. The reduction of variability in the distribution of the fatigue-life data of SCC and SCFRC compared to NVC and NVFRC may be attributed to the relatively denser and more homogeneous composition of SCC. Without consolidation, the influence of intrinsic deficiencies and material defects due to bleeding or segregation induced by improper vibration practice may be avoided. As a result, the homogeneity of SCC can be ensured and may substantially enhance the mechanical properties and reliability of structural members. In addition, the alignment of the steel fibers in the direction of flow, as reported by few researchers,13,21 may also be attributed to the better flexural fatigue properties of SCFRC compared to NVFRC.

Goodness-of-fit testAs shown in the preceding section, the fatigue-life distri-

butions of SCC and SCFRC at different stress levels can approximately be described by the two-parameter Weibull distribution. Further, the values of the correlation coefficient at each stress level also substantiated this. In addition, the Kolmogorov-Smirnov test was applied as goodness of fit to the fatigue-life data at each stress level tested in this inves-tigation; it was observed that the model was acceptable at a 5% significance level.29,30,32 The calculations to this effect are not given.

Flexural fatigue performance of SCC and SCFRCIn the preceding sections, the flexural fatigue-life data

of SCC and SCFRC were shown to follow the two-param-eter Weibull distribution at different stress levels. This can further be used to calculate the fatigue lives corresponding to different failure probabilities Pf.29

Substituting 1– Pf = LN in Eq. (3), the following relation is obtained

( )1ln ln ln ln( )1 f

n uP

= α − α −

(11)

Rearranging Eq. (11)

1

1ln ln ln( )1

lnf

nP

N −

+ α − = α

(12)

Equation (12) can be used to calculate the fatigue life N for a particular probability of failure Pf. Using the average values of the parameters of the Weibull distribution for fatigue-life data at a given stress level S as listed in Table 5, Eq. (12) is used to calculate the fatigue lives for SCC and SCFRC with 0.5, 1.0, and 1.5% of volume fractions of steel fibers, corresponding to a failure probability of 10%—that is, Pf = 0.1. These calculated values of fatigue lives are plotted in Fig. 8 to obtain fatigue curves for SCC and SCFRC with 0.5, 1.0, and 1.5% of volume fraction of steel fibers for a failure probability of 0.1.

Fig. 8—Fatigue curves of SCC and SCFRC corresponding to 10% probability of failure (Pf = 0.1) using single-log failure equation.


In this study, a single-log fatigue equation, Eq. (1), was used to evaluate the performance of SCC and SCFRC corre-sponding to a probability of failure Pf = 0.1. The single-log fatigue equation is commonly used by researchers to describe the relation between stress level S and fatigue life N.22,25,37 In Eq. (1), the fatigue performance is dependent on the two important coefficients/parameters a and b. The parameter a reflects the height of the fatigue curve. The larger the param-eter a, the higher the fatigue curve. The parameter b reflects the steep degree of the fatigue curve. The larger the parameter b, the steeper the fatigue curve, and the fatigue life of concrete is more sensitive to the change in stress.37 The parameters a and b of Eq. (1) are obtained from regression analysis for the fatigue curves of SCC and SCFRC plotted in Fig. 8. The estimated fatigue equations and the corresponding correla-tion coefficients are also presented in Table 6. The param-eters a and b of the single-log equation and their enhanced extent generated for SCC and SCFRC with 0.5, 1.0, and 1.5% of volume fraction of steel fibers for a failure probability of 10%—that is, Pf = 0.1—are also listed in Table 6. It can be seen that the regression parameters of fatigue equations of all SCFRC mixtures are increased to a different extent, and all correlation coefficients are higher than 0.97. The increase in the value of a indicates that the flexural fatigue performance of all SCFRC mixtures is significantly improved. The value of b increased for all SCFRC mixtures, indicating an increase in the sensitivity of their fatigue lives to change of stress. Compared with the SCC and other SCFRC mixtures, SCFRC with 1.0% of steel fibers has the largest enhanced extent of 7.30% for a and the smallest enhanced extent of 0.78% for b, thus indicating that the fatigue performance of SCFRC with 1.0% of steel fibers is improved to the largest extent and the sensitivity of its fatigue life to change in stress is increased to the smallest extent.

Table 7 shows the theoretic fatigue lives of various SCC and SCFRC mixtures calculated by a single-log fatigue equation corresponding to a 10% probability of failure at

five different stress levels. It can be seen that the theoretic fatigue lives of SCFRC containing 0.5, 1.0, and 1.5% steel fibers increased to a different extent. With the increasing stress level, the enhanced extent of the theoretic fatigue lives of the SCFRC mixtures increased, which indicates that the SCC containing steel fibers—that is, SCFRC—has an excel-lent fatigue performance, particularly at higher stress levels (corresponding to heavy traffic load) compared with SCC.

It can be seen from Table 7 that the enhanced extent of the theoretic fatigue life of SCFRC with 1.0% of steel fibers is the highest—that is, 908%—at a stress level of 0.90. A similar increase was observed for SCFRC with 1.0% of steel fibers at all the other stress levels, indicating a superior flex-ural fatigue performance of SCFRC1.0 compared to SCC and other SCFRC mixtures used in the investigation. It can also be seen from the results that the fatigue performance of SCC is improved with the addition of steel fibers. It may, however, be noted that the fatigue performance herein is represented in terms of applied maximum fatigue stress expressed as a percentage of corresponding static flexural stress—that is, in terms of stress level S. In contrast, when the fatigue perfor-mance is examined in terms of applied maximum fatigue stress fmax, the ranking differs. Increasing the fiber content from 0.5 to 1.5% seems to improve the fatigue performance in terms of applied maximum fatigue stress. Similar trends were obtained by earlier investigators38,39 while studying the flexural fatigue performance of NVFRC.

It may be noted that the results, such as parameters of the Weibull distribution and theoretic fatigue life of SCFRC reported in this paper, are applicable to the type, size, and volume fraction of the fibers used; therefore, additional work is required to generate more data for other types, sizes, and volume fraction of fibers.

CONCLUSIONS1. The probability distributions of fatigue life of SCC

and SCFRC, at different stress levels, can be approximately

Table 6—Single-log fatigue equation and its coefficients a and b corresponding to 10% probability of failure (Pf = 0.1) for SCC and SCFRC

Mixture Fatigue equation, S = a – blog(N) Cc* a Enhanced extent, % b Enhanced extent, %

SCC S = 1.0912 – 0.0762log(N) 0.994 1.0912 0 0.0762 0

SCFRC0.5 S = 1.1532 – 0.0836log(N) 0.999 1.1532 5.57 0.0836 9.71

SCFRC1.0 S = 1.1698 – 0.0768log(N) 0.996 1.1698 7.20 0.0768 0.79

SCFRC1.5 S = 1.1397 – 0.0788log(N) 0.978 1.1397 4.44 0.0788 3.41*Cc is correlation coefficient.

Table 7—Theoretic fatigue lives (number of cycles) for SCC and SCFRC calculated by single-log fatigue equation corresponding to 10% probability of failure

Stress level, S → 0.90 0.85 0.80 0.75 0.70

Mixture ↓

SCC Theoretic fatigue life 323 1463 6630 30,039 136,100

SCFRC0.5Theoretic fatigue life 1068 4234 16,783 66,522 263,665

Enhanced extent, % 230 189 153 121 93

SCFRC1.0Theoretic fatigue life 3259 14,590 65,329 292,513 1,309,747

Enhanced extent, % 908 897 885 873 862

SCFRC1.5Theoretic fatigue life 1101 4747 20,460 88,192 380,145

Enhanced extent, % 240 224 208 193 179


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modeled by the two-parameter Weibull distribution with a statistical correlation coefficient greater than 0.90.

2. Higher values of shape parameters for SCC as well as SCFRC shows a reduction in the variability in the distribu-tion of fatigue life of SCC and SCFRC compared to NVC and NVFRC, respectively.

3. The lower values of the shape parameter for SCFRC, as compared with SCC, show that the variability in the distri-bution of the fatigue life of SCFRC is larger and further increases for lower fatigue stress levels.

4. Failure probability was incorporated in the single-log equation to examine the flexural fatigue performance of SCC and SCFRC. Theoretic fatigue lives and curves were gener-ated for SCC and SCFRC, corresponding to a failure prob-ability of 10%—that is, 0.1.

5. The enhanced extent of the theoretic fatigue lives of SCFRC increased to a different extent with an increase in fiber content, thereby depicting a better flexural fatigue performance of SCFRC as compared to SCC—particularly at higher stress levels.

NOTATIONa, b = regression parameters of single-log fatigue equationfmax = maximum fatigue stressfmin = minimum fatigue stressfr = static flexural strengthLN = survival probabilityN = number of cycles to failurePf = probability of failureR = stress ratio = fmin/fmax

S = stress level = fmax/fr

u = scale parametera = shape parameterG() = gamma functionm = mean of data samples = standard deviation of data sample under consideration

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Specification, Production and Use, Experts for Specialised Construction and Concrete Systems, Farnham, UK, 2005, 68 pp.

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In ACI STRUCTURAL JoURnALThe American Concrete Institute also publishes the ACI Structural Journal. This section presents brief synopses of papers appearing in the current issue.

From the September-October 2012 issue

PDF versions of these papers are available for download at the ACI website, www.concrete.org, for a nominal fee.

and errors in experimental works and increase the credibility of test results. The authors analyzed the accuracy and reliability of various measurement alternatives employed in this study. Acceptable performance was demonstrated when electrical resistance strain gauges (ERSGs) were placed on strands’ surfaces at intervals of 5.9 in. (150 mm) and on concrete surfaces with cover thicknesses not greater than 3.0 in. (75 mm). Also, strain readings from gauges mounted on strands can be used to estimate the amount of prestress through an adjustment process. High-temperature steam curing somewhat adversely affects transfer length; and when strand is debonded near the end as is common in the fabrication of precast prestressed concrete, a significant reduction of transfer length may occur, especially at the cut end. The application of a sudden detensioning method by disc cutting produces different transfer lengths at each cut and dead end.

109-S55—Compatibility Strut-and-Tie Modeling: Part I—Formulationby Reece M. Scott, John B. Mander, and Joseph M. Bracci

A compatibility-based strut-and-tie model (C-STM) intended for analyzing the nonlinear force-deformation behavior of disturbed regions and structural concrete deep beams is presented. In addition to the normal strut-and-tie force equilibrium requirements, the proposed C-STM accounts for nonlinear behavior using nonlinear constitutive relations for cracked reinforced concrete. The model is implemented using commercially available structural analysis software, SAP2000TM. To assess C-STM accuracy, convergence studies using different truss representations are explored. Particular emphasis is placed on investigating the behavior of deep cantilevered beams to provide insight into the progression of nonlinear response leading to the ultimate shear failure. New developments for modeling the nonlinear behavior of concrete compression chord members and compression-softening effects of diagonal concrete struts are proposed. The implementation is presented in the companion paper.

109-S56—Compatibility Strut-and-Tie Modeling: Part II—Implementationby Reece M. Scott, John B. Mander, and Joseph M. Bracci

This paper presents the implementation and computational validation of a compatibility-based strut-and-tie model (C-STM) presented in a companion paper intended for analyzing the nonlinear force-deformation behavior of disturbed regions and structural concrete deep beams. The C-STM is used to predict the force-deformation response and internal nonlinear strain behavior of previously conducted large-scale reinforced concrete bridge bent-cap experiments. Additionally, the experimental results are compared with current code-based approaches to illustrate how the C-STM can be used as a minimalist computational analysis tool to provide an accurate prediction of the structure’s force-displacement response. A comprehensive description of how the C-STM is implemented into structural analysis software SAP2000TM is given to provide a step-by-step modeling procedure.

109-S57—Development Length of Unconfined Conventional and High-Strength Steel Reinforcing Barsby Amr Hosny, Hatem M. Seliem, Sami H. Rizkalla, and Paul Zia

The development length equation specified by ACI 318-08 and the similar equation recommended by ACI 408R-03 are based on extensive test results using conventional reinforcement conforming to ASTM A615/A615M and ASTM A706/A706M. With the development of new ASTM A1035/A1035M high-strength steel reinforcement, several studies have been conducted to examine whether the current equations are applicable for the new high-strength reinforcing steel. These studies have shown that the current equations could, in some cases, overestimate the bond strength of high-strength steel bars. This paper proposes a new equation for the bond strength of unconfined reinforcing

109-S51—Effect of Load Distribution and Variable Depth on Shear Resistance of Slender Beams without Stirrupsby Alejandro Pérez Caldentey, Patricio Padilla, Aurelio Muttoni, and Miguel Fernández Ruiz

The shear resistance of elements without stirrups has mainly been investigated by test setups involving simply supported beams of constant thickness subjected to one- or two-point loading, and most of the formulas included in codes have been adjusted using this experimental background. Most design situations, however, involve constant or triangular distributed loading on tapered members. Also, there seems to be few shear tests involving cantilever structures subjected to distributed loading. These structures fail in shear near the clamped end.

This paper presents a specific testing program. It investigates the influence of load distribution and tapered beam geometrics on shear strength. The experimental program consists of eight slender beams without stirrups. The setup allowed direct comparisons between point and distributed loading. The experimental results showed a significant influence of the type of loading and of tapered geometries on the shear strength. Based on these results, and using the fundamentals of the critical shear crack theory, a consistent physical explanation of the observed failure modes and differences in shear strength is provided. Also, comparisons to current design provisions (ACI 318-08 and EC2) are discussed.

109-S52—Optimal Strut-and-Tie Models Using Full Homogenization Optimization Methodby Juan Pablo Herranz, Hernán Santa María, Sergio Gutiérrez, and Rafael Riddell

An optimization method based on homogenization is proposed for finding optimal strut-and-tie (ST) models for reinforced concrete (RC) elements. The method uses a layout that minimizes displacement for a given loading state in a linearly elastic regime by mixing two materials. Although this optimal layout might contain fine mixtures, one can still obtain a strongly resembling ST model without mixtures that performs closely to the optimal configuration through a penalization procedure. Two examples from the ST literature are used to illustrate the application of the method: the dapped beam and the beam on beam. The reinforcement layouts obtained make the element more efficient in terms of ultimate load divided by the weight of the steel used and having smaller deflections and crack widths.

109-S53—Cyclic Load Testing for Integrity Evaluation of Prestressed Concrete Girdersby Francisco Barrios and Paul H. Ziehl

This study focuses on the performance and validation of the 24-hour load test (24h LT) method and the cyclic load test (CLT) method as applied to full-scale lightweight and normalweight self-consolidating prestressed concrete girders. It examines data obtained from the four-point loading tests of six full-scale T-girders and applies the current criteria from these methodologies to evaluate the presence of damage and structural integrity. The experimental results indicate that the recovery criteria of the 24h LT method were insensitive to damage and hence did not provide a satisfactory integrity assessment of the members. The permanency and repeatability criteria of the CLT were insensitive to damage for the girders studied. The global integrity parameter based on the deviation from linearity criterion from the CLT is proposed for the quantitative assessment of the level of damage in prestressed concrete girders; the results indicate good correlation with the experimental data.

109-S54—Methodological Aspects in Measurement of Strand Transfer Length in Pretensioned Concreteby Ho Park, Zia Ud Din, and Jae-Yeol Cho

This study assessed experimental methodological factors that might affect the estimation of transfer length of pretensioned concrete to minimize trials


bars for all three types of steel. The proposed equation for high-strength steel is compared to extensive test data reported in the literature and is found to be more accurate than ACI 318-08 and ACI 408R-03 equations specified for conventional reinforcement.

109-S58—Improved Algorithm for Efficient and Realistic Creep Analysis of Large Creep-Sensitive Concrete Structuresby Qiang Yu, Zdenek P. Bažant, and Roman Wendner

A recent compilation of data on numerous large-span prestressed segmentally erected box girder bridges revealed gross underestimation of their multi-decade deflections. The main cause was identified as incorrect and obsolete creep prediction models in various existing standard recommendations and is being addressed in a separate study. However, previous analyses of the excessive deflections of the Koror-Babeldaob (KB) Bridge in Palau and of four Japanese bridges have shown that a more accurate method of multi-decade creep analysis is required. This paper provides a systematic and comprehensive presentation, appropriate not only for bridges but also for any large creep-sensitive structure. For each time step, the solution is reduced to an elastic structural analysis with generally orthotropic elastic moduli and eigenstrains. This analysis should normally be three-dimensional. It can be accomplished with a commercial finite element code such as ABAQUS. Based on the Kelvin chain model, the integral-type creep law is converted to a rate-type form with internal variables. For time steps short enough to render aging during each step to be negligible, a unique continuous retardation spectrum for each step is obtained by Laplace transform inversion using simple Widder’s formula. Discretization of the spectrum then yields the current Kelvin chain moduli. The rate-type creep analysis is computationally more efficient than the classical integral-type analysis. Also, it is possible to take into account the evolution of various inelastic and nonlinear phenomena. Finally, the advantages compared to the existing commercial programs are pointed out and illustrated by a simple example.

109-S59—Design of Thick Concrete Plates Using Strut-and-Tie Modelby E. Rizk, H. Marzouk, and R. Tiller

A strut-and-tie model (STM) is developed to model the punching shear behavior of thick concrete plates. This model provides a quick and simple approach to punching shear behavior. It is applicable to both normal- and high-strength concrete under symmetric and nonsymmetric loading with and without shear reinforcement. The STM for symmetric punching consists of a “bottle-shaped” compressive zone in the upper section of the slab depth, leading to a “rectangular-stress” compressive zone in the lower section depth. An equation based on failure criteria for the STM is used to model the behavior in the lower compressive stress zone. Another STM is also developed to rationally model nonsymmetric punching shear behavior due to unbalanced moment transfer and symmetric punching behavior of concrete slabs with shear reinforcement. The results of the STMs for symmetric and nonsymmetric loading with and without shear reinforcement were compared to experimental test results performed and published by others. The results of the STMs showed excellent agreement with available test results.

109-S60—Influence of Span-Depth Ratio on Behavior of Externally Prestressed Concrete Beamsby T. J. Lou, A. V. Lopes, and S. M. R. Lopes

This paper describes a numerical study on the flexural behavior of concrete beams prestressed with external tendons, focusing on the effect of the span-depth ratio (L/dp) on the response characteristics and the ultimate stress in tendons and ductility behavior. A nonlinear model, calibrated by the available experimental results of externally prestressed specimens, is used for the parametric evaluation. The results show that the second-order effects of externally prestressed beams with two deviators at third points become increasingly important with the increase of the L/dp and that, due to these effects, a higher L/dp registers a lower ultimate moment capacity. The effect of the L/dp on the ultimate tendon stress is dependent on the type of loading for internal unbonded tendons and on the configuration of deviators for external tendons. Also, irrespective of second-order effects, the L/dp has an insignificant effect on the ductility of beams, and center-point loading mobilizes higher deflection ductility than third-point or uniform loading.

109-S61—Life-Cycle Cost Analysis of Carbon Fiber-Reinforced Polymer Reinforced Concrete Bridgesby Nabil F. Grace, Elin A. Jensen, Christopher D. Eamon, and Xiuwei Shi

This paper presents a life-cycle cost analysis (LCCA) of prestressed concrete highway bridges using carbon fiber-reinforced polymer (CFRP) reinforcement bars and strands. Side-by-side box beam and AASHTO beam bridge structures were considered over several span lengths and traffic volumes. The results show that despite the higher initial construction cost of CFRP reinforced bridges, they can be cost-effective when compared to traditional steel-reinforced bridges. The most cost-efficient design was found to be a medium-span CFRP reinforced AASHTO beam bridge located in a high-traffic area. A probabilistic analysis revealed that there is greater than a 95% probability that the CFRP reinforced bridge will become the least expensive option between 20 and 40 years of service, depending on traffic volume and bridge geometry. The break-even year between CFRP and steel reinforcement is typically at the time of the first major repair activity on the steel-reinforced concrete bridge.

109-S62—Shear Capacity Prediction of Reinforced Concrete Beams without Stirrups Using Fracture Mechanics Approachby Shilang Xu, Xiufang Zhang, and Hans W. Reinhardt

This study presents an analytical shear strength prediction equation for lightly reinforced slender concrete beams without stirrups based on the phenomenological experiment observations. The concept of the loss of bond performance between concrete and longitudinal reinforcement was used in these beams to explain the potential cause for the sudden release of longitudinal reinforcement from wrapping concrete. In the proposed equation, shear capacity was related to bond fracture resistance by introducing a new parameter: Mode II fracture toughness KIIc. The equation showed the size effect with effective depth to the power of –1/2 and was evaluated using test data published in other sources. Comparisons between the proposed formula and other prediction equations indicated that, for lightly reinforced slender concrete beams without stirrups, this developed formula can estimate the shear strength of beams with varying concrete strengths, shear span-depth ratios (as /d), longitudinal reinforcement ratios, and beam depths with reasonable accuracy.

109-S63—Repair of Corroded Prestressed Concrete Piles of Harbor Landing Stages by Tseng-Cheng Lin, Chyuan-Hwan Jeng, Chung-Yue Wang, and Ting-Hung Jou

Corrosion and deterioration of concrete piles are common problems for wharf structures. In Taiwan, corroded prestressed concrete (PC) piles supporting landing stage structures are frequently repaired and strengthened using carbon fiber-reinforced polymer (CFRP) jacketing. This paper presents an investigation on corroded and CFRP-repaired PC piles for landing stage structures. Seven reduced-scale PC pile specimens were tested to investigate their cyclic lateral load-carrying behavior. The experimental results showed that the CFRP jacketing is rather efficient in terms of lateral strength, capacity for deformation and energy dissipation, and stiffness degradation, and is effective in repairing the seismic-resistance capacity of the corroded specimens. This study also conducted finite element (FE) analyses to analyze the tests of the seven specimens, achieving good corroboration between the analytical results and the tests.

109-S64—Simplified Method for Nonlinear Dynamic Analysis of Shear-Critical Framesby Serhan Guner and Frank J. Vecchio

A nonlinear static analysis method was recently developed for the performance assessment of plane frames. This method’s primary advantage is its ability to accurately represent shear effects coupled with axial and flexural behaviors through a simple modeling approach suitable for large-scale applications. This study further develops this method to enable a dynamic load analysis capability under impact, blast, and seismic loads. Newly developed and implemented formulations are presented. The method is applied to 11 previously tested specimens, subjected to impact and seismic loads, to examine its accuracy, reliability, and practicality. The method is found to simulate the overall experimental behaviors with a high degree of accuracy. Strengths, peak displacements, stiffnesses, damage, and failure modes (including shear-critical behaviors) and vibrational characteristics are calculated accurately. The method provides unconditional numerical stability and requires a fraction of the computation time demanded by micro finite element methods.

ACIMATERIALSJOURNAL

This journal and a companion periodical, ACI Structural Journal, continue the publishing tradition the Institute started in 1904. Information published in ACI Structural Journal includes: structural design and analysis of concrete elements and structures, research related to concrete elements and structures, design and analysis theory, and related ACI standards and committee reports.

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