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Suggested Citation:"Chapter 2 - Literature Review." National Academies of Sciences, Engineering, and Medicine. 2008. Transfer, Development, and Splice Length for Strand/Reinforcement in High-Strength Concrete. Washington, DC: The National Academies Press. doi: 10.17226/13916.
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Suggested Citation:"Chapter 2 - Literature Review." National Academies of Sciences, Engineering, and Medicine. 2008. Transfer, Development, and Splice Length for Strand/Reinforcement in High-Strength Concrete. Washington, DC: The National Academies Press. doi: 10.17226/13916.
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Suggested Citation:"Chapter 2 - Literature Review." National Academies of Sciences, Engineering, and Medicine. 2008. Transfer, Development, and Splice Length for Strand/Reinforcement in High-Strength Concrete. Washington, DC: The National Academies Press. doi: 10.17226/13916.
×
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Suggested Citation:"Chapter 2 - Literature Review." National Academies of Sciences, Engineering, and Medicine. 2008. Transfer, Development, and Splice Length for Strand/Reinforcement in High-Strength Concrete. Washington, DC: The National Academies Press. doi: 10.17226/13916.
×
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Suggested Citation:"Chapter 2 - Literature Review." National Academies of Sciences, Engineering, and Medicine. 2008. Transfer, Development, and Splice Length for Strand/Reinforcement in High-Strength Concrete. Washington, DC: The National Academies Press. doi: 10.17226/13916.
×
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Suggested Citation:"Chapter 2 - Literature Review." National Academies of Sciences, Engineering, and Medicine. 2008. Transfer, Development, and Splice Length for Strand/Reinforcement in High-Strength Concrete. Washington, DC: The National Academies Press. doi: 10.17226/13916.
×
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Suggested Citation:"Chapter 2 - Literature Review." National Academies of Sciences, Engineering, and Medicine. 2008. Transfer, Development, and Splice Length for Strand/Reinforcement in High-Strength Concrete. Washington, DC: The National Academies Press. doi: 10.17226/13916.
×
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Suggested Citation:"Chapter 2 - Literature Review." National Academies of Sciences, Engineering, and Medicine. 2008. Transfer, Development, and Splice Length for Strand/Reinforcement in High-Strength Concrete. Washington, DC: The National Academies Press. doi: 10.17226/13916.
×
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Suggested Citation:"Chapter 2 - Literature Review." National Academies of Sciences, Engineering, and Medicine. 2008. Transfer, Development, and Splice Length for Strand/Reinforcement in High-Strength Concrete. Washington, DC: The National Academies Press. doi: 10.17226/13916.
×
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Suggested Citation:"Chapter 2 - Literature Review." National Academies of Sciences, Engineering, and Medicine. 2008. Transfer, Development, and Splice Length for Strand/Reinforcement in High-Strength Concrete. Washington, DC: The National Academies Press. doi: 10.17226/13916.
×
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Suggested Citation:"Chapter 2 - Literature Review." National Academies of Sciences, Engineering, and Medicine. 2008. Transfer, Development, and Splice Length for Strand/Reinforcement in High-Strength Concrete. Washington, DC: The National Academies Press. doi: 10.17226/13916.
×
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Suggested Citation:"Chapter 2 - Literature Review." National Academies of Sciences, Engineering, and Medicine. 2008. Transfer, Development, and Splice Length for Strand/Reinforcement in High-Strength Concrete. Washington, DC: The National Academies Press. doi: 10.17226/13916.
×
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Suggested Citation:"Chapter 2 - Literature Review." National Academies of Sciences, Engineering, and Medicine. 2008. Transfer, Development, and Splice Length for Strand/Reinforcement in High-Strength Concrete. Washington, DC: The National Academies Press. doi: 10.17226/13916.
×
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Suggested Citation:"Chapter 2 - Literature Review." National Academies of Sciences, Engineering, and Medicine. 2008. Transfer, Development, and Splice Length for Strand/Reinforcement in High-Strength Concrete. Washington, DC: The National Academies Press. doi: 10.17226/13916.
×
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Suggested Citation:"Chapter 2 - Literature Review." National Academies of Sciences, Engineering, and Medicine. 2008. Transfer, Development, and Splice Length for Strand/Reinforcement in High-Strength Concrete. Washington, DC: The National Academies Press. doi: 10.17226/13916.
×
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Suggested Citation:"Chapter 2 - Literature Review." National Academies of Sciences, Engineering, and Medicine. 2008. Transfer, Development, and Splice Length for Strand/Reinforcement in High-Strength Concrete. Washington, DC: The National Academies Press. doi: 10.17226/13916.
×
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Suggested Citation:"Chapter 2 - Literature Review." National Academies of Sciences, Engineering, and Medicine. 2008. Transfer, Development, and Splice Length for Strand/Reinforcement in High-Strength Concrete. Washington, DC: The National Academies Press. doi: 10.17226/13916.
×
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Suggested Citation:"Chapter 2 - Literature Review." National Academies of Sciences, Engineering, and Medicine. 2008. Transfer, Development, and Splice Length for Strand/Reinforcement in High-Strength Concrete. Washington, DC: The National Academies Press. doi: 10.17226/13916.
×
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Suggested Citation:"Chapter 2 - Literature Review." National Academies of Sciences, Engineering, and Medicine. 2008. Transfer, Development, and Splice Length for Strand/Reinforcement in High-Strength Concrete. Washington, DC: The National Academies Press. doi: 10.17226/13916.
×
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Suggested Citation:"Chapter 2 - Literature Review." National Academies of Sciences, Engineering, and Medicine. 2008. Transfer, Development, and Splice Length for Strand/Reinforcement in High-Strength Concrete. Washington, DC: The National Academies Press. doi: 10.17226/13916.
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62.1 Introduction The comprehensive and critical literature review under- taken during NCHRP Project 12-60 is described in this chap- ter. In this report, important findings from prior research are reviewed, with particular attention to the impact of these findings on the work plan for NCHRP Project 12-60. The objective of the work related to prestressing strand was to gather and synthesize existing data and information on the transfer length and development length of strand with diam- eters up to 0.6 in. In the area of mild reinforcement, the effort concentrated on development and splice length in tension of individual bars and development length of bars in tension an- chored with standard hooks. The database constructed from this effort includes 71 tension development and splice tests of specimens with top cast uncoated reinforcing bars, 493 speci- mens with bottom cast uncoated reinforcing bars, 27 speci- mens with top cast epoxy-coated bars, and 48 with bottom cast epoxy-coated bars. In addition, 33 specimens with un- coated bars terminated with standard hooks and 13 specimens reinforced with epoxy-coated bars have been reviewed. A comprehensive analysis of the data collected was con- ducted to identify issues and needs related to bond of strand and mild steel in high-strength concrete. This analysis assisted in the identification of several key variables that are likely to affect the transfer and development of prestressing strands, development/splice length of bars in tension, and develop- ment length of bars in tension terminated with standard hooks. Some of these variables are currently included in the AASHTO LRFD Bridge Design Specifications while some are not. In the area of transfer length and development length of prestressing strand, specifications do not account for variables such as concrete strength, strand size, “top bar” effects, epoxy coating, bond quality of individual strand samples, and struc- tural behavior issues (e.g., the interaction of shear and bond). The work plan for NCHRP Project 12-60 included procedures and testing to evaluate some, but not all, of these effects. Other issues that may influence the transfer and development of pre- stressing strands include confining reinforcement and strand spacing. The research reported in the literature indicates that 0.6-in. strand can be spaced at 2.0 in. center to center (c/c) or that 0.5-in. strands may be spaced at 1.75 in. without penalty to the transfer and development of strands. The reported research also indicates that confining reinforcement has little or no effect on transfer length of strands, but it can be quite beneficial to strand development. Standardized confinement details were employed in beam testing where warranted. In the area of bond of mild reinforcement, the single most important issue not currently accounted for in the AASHTO LRFD Bridge Design Specifications is the effect of confining reinforcement on the bond strength of tension reinforcement in the case of splitting type failures. This parameter is espe- cially important as bars are being developed in higher strength concretes. The review also revealed that for epoxy-coated bar development/splice length and development length of bars terminated with standard hooks there is a paucity of data on concretes with cylinder strengths above 10 ksi. A significant effort during the initial 6 months of the NCHRP Project 12-60 study was focused on identifying and evaluating testing protocols related to the experimental work to be conducted. In the area of bond in concrete of prestress- ing strand, particular attention was given to the surface char- acterization tests to evaluate strand “bond-ability.” There are three tests that have been offered in recent years as possible tests to standardize acceptance of strand based on its “bond- ability”: (1) the Moustafa Test, where untensioned strands are pulled from large concrete blocks; (2) the PTI Bond Test, where untensioned strands are pulled from a neat cement mortar; and (3) the NASP Bond Test, where untensioned strands are pulled from a sand cement mortar. In testing per- formed by the North American Strand Producers (NASP), the NASP Bond Test has proven to be the most reliable test of the three. It has produced test results from “blind trials” with the best repeatability and reproducibility. C H A P T E R 2 Literature Review

7In three separate rounds of testing, the Moustafa Test (now called the Large Block Pull-Out Test) was performed at differ- ent sites to determine its reproducibility across sites. In NASP Round I testing, the Moustafa Test was performed at Coreslab Structures in Colorado and at Florida Wire and Cable Co. (FWC). For the purpose of carrying out Moustafa Tests, FWC built a completely automated testing machine so that the Moustafa procedures could be precisely followed. Round I test- ing showed widely dissimilar results from the two testing sites. In NASP Round II, the Moustafa Test and the PTI Bond Tests were performed at three testing sites: Coreslab Structures, FWC, and the University of Oklahoma. Additionally, the NASP Bond Test was introduced in an early form as a test very similar to the PTI Bond Test except that a sand-cement mortar was used. Seven different strand samples were shipped to the different testing sites. The trials were blind. Again, the Moustafa Test failed to produce reproducible results across testing sites. Of the three tests, the NASP Bond Test showed the highest statistical correlation across testing sites. In the NASP Round III testing, a more refined version of the NASP Bond Test again outper- formed the Moustafa Test in blind trials at the three testing sites. In all three rounds of testing, when the Moustafa Test was used, it failed to produce results that were consistent across sites. The NASP Bond Test proved more reliable at providing the same or similar results across testing sites in Rounds II, III, and IV. Be- cause of the NASP Bond Test’s more consistent results, the NCHRP Project 12-60 testing program was built upon the NASP Bond Test. The review conducted on testing for development/splice length of deformed bars in tension showed that the generally recommended testing protocol for full-scale specimens because of both the relative ease of fabrication and the realistic state of stress achieved during testing is the beam-splice specimen. Thus, beam splice specimens were used in the development of experimental data related to development/splice length of mild reinforcement during the course of the NCHRP Project 12-60 study. It is well established that testing protocols to evaluate de- velopment and splice length requirements for deformed bars and wire in tension must be of an appropriate scale, containing more than one bar or wire, with due regard for a realistic trans- fer of force between concrete and steel reinforcement and cover/bar spacing effects. Splice tests have in the past been ac- curate simulations of real conditions in structures; however, de- velopment length tests have been largely conducted using pull- out tests, in which splitting failures are purposely avoided. As a result, the bond stresses developed along splices are low com- pared with the bond stresses along a bar in a pull-out test. This difference in test methods is responsible for large differences in code-required anchorage lengths for splices and development of single bars. Pull-out failures occur in cases of high confine- ment and short bonded lengths. In most structural applications, however, splitting failures tend to control. Beam-splice speci- mens are deemed to represent larger-scale specimens designed to directly measure development and splice strength in full- scale members. The experimental work supporting the current require- ments in the AASHTO LRFD Bridge Design Specifications and ACI 318-05: Building Code Requirements for Structural Concrete and Commentary (2005) for development of stan- dard hooks in tension was conducted using a test setup rep- resenting an exterior beam column joint. Because of the paucity of data on concrete strengths above 10 ksi, the evalu- ation of uncoated and epoxy-coated bars terminated with standard hooks in tension to normal-weight concrete with compressive strength up to 15 ksi was performed using a sim- ilar exterior beam column joint test setup. The results of the initial work of NCHRP Project 12-60 confirmed the basic premises stated in the original project proposal. Thus, the efforts of the experimental program and the order of priority of these efforts remained as originally stated. The experimental program focused on the following major efforts listed in priority order: 1. Determining design equations for transfer and develop- ment length of strand in prestressed concrete bridge mem- bers. Variables included concrete strength at release, con- crete strength at time of development length testing, use of air-entraining admixtures, “top bar effects,” and strand size. 2. Development and splice length in tension of reinforcing bars. Variables included concrete strength, bar size, con- crete cover/bar spacing, amount of transverse reinforce- ment, epoxy coating, and casting position. 3. Development length in tension of bars terminated with standard hooks. Variables included concrete strength, bar size, concrete cover/bar spacing, amount of transverse reinforcement, and epoxy coating. A comprehensive article-by-article review of Section 5 of the 2nd edition of the AASHTO LRFD Bridge Design Specifi- cations with the 1999, 2000, and 2001 interim revisions (AASHTO 1998) was conducted during the initial 6 months of the NCHRP Project 12-60 study. In this review, the provi- sions of Section 5 that directly or indirectly affect transfer and development length of prestressing strand and develop- ment/splice length of mild reinforcement by the use of high- strength concrete were extracted and critically reviewed. 2.2 Literature Review 2.2.1 Strand Transfer and Development Length A number of experimental investigations related to high- strength concrete have been conducted in North America and overseas. Hence, a significant body of knowledge currently ex- ists with respect to the performance of high-strength concrete

8members. Amongst the experimental data are various studies dealing with transfer length and development length of pre- stressing strand and splice length and development length of black and epoxy-coated reinforcement. In this study, a com- prehensive and critical literature review was undertaken to gather and synthesize existing data and information related to the transfer length and development length of prestressing strand with diameters up to 0.6 in., and development and splice length in tension and compression of individual bars, bundled bars, and welded wire reinforcement and develop- ment length of standard. The literature review centered on collecting information on testing protocols for determining surface bond character- istics of strand, performance of members containing trans- verse reinforcement, bond and transfer length, and tests addressing deformation capacity. Information available from the field—including FHWA showcase projects and the unpublished experiences of engineers, bridge owners, and producers—was reviewed and used to supplement other work conducted in this study. The development of reliable code expressions for transfer and development of prestressing strand is made more diffi- cult by the large experimental scatter reported by researchers over the past 40 years. The original code expressions for transfer and development length of pretensioned strands were developed from testing performed in the late 1950s and early 1960s on Grade 250, stress-relieved strand (Hanson and Kaar 1959; Kaar, LaFraugh, and Mass 1963; Tabatabai and Dickson 1993). Based on these early tests, the ACI Building Code (ACI 2005) and the AASHTO LRFD Bridge Design Specifications adopted provisions governing the design for strand transfer and development. Manufacturing innovation has brought about Grade 270 low-relaxation strand as the in- dustry standard, while the code expressions for transfer and development length have changed very little. Furthermore, contemporary strand production employs induction heating to stress relieve strand, whereas convection heating was used in the late 1950s and early 1960s. Convec- tion heating created hotter surface temperatures on strand that may have burned off much of the surface residues remaining from the wire drawing process. Today’s processes, using induction heating, may have created surface tempera- tures lower than those created by convection heating and thereby may have effectively changed the bonding character- istics of the surface of prestressing strands (Rose and Russell 1997). In the mid-1980s, Cousins, Johnston, and Zia (1990) measured transfer lengths that exceeded the standard design predictions by a wide margin. Their findings led FHWA to adopt a moratorium on the use of 0.6 in. diameter strands and to increase the development length for other sizes of pre- stressing strands. The FHWA action led to the creation of a large number of research programs intent on measuring the transfer and development of prestressing strands. Research was performed at the University of Texas (Russell and Burns 1996, 1997), Florida DOT (Shahawy, Issa, and Batchelor 1992), McGill University (Mitchell et al. 1993), and Auburn University (Cousins et al. 1993). The arbitrary 1.6 multiplier from the original FHWA moratorium is now incorporated into the AASHTO LFRD Bridge Design Specifications. By the mid-1990s, it became apparent that the studies ex- amining the transfer and development of prestressing strands had not resulted in a consensus on design standards. As a whole, the research displayed a large scatter of the test results, with measured transfer lengths for 0.5-in. strand ranging from a low of less than 20 in. to a high of more than 60 in. Thus, it became apparent that other variables were in play and that such variables were not properly accounted for in either design equations or specifications. Since the mid-1990s, research work has concentrated on developing a standardized test to assess the bond characteris- tics of individual prestressing strands. It was suspected that different strand manufacturers produced strand with quite dissimilar bonding characteristics. Hence, it was important to quantify the bonding characteristics of an individual strand before the transfer length and development length data would be meaningful. To that end, three or four different testing programs were undertaken to assess the viability of various “standardized tests” and the suitability of such tests for predicting the “bond-ability” of prestressing strand. The first such testing program was developed by Rose and Russell (1997). The various testing programs found little correlation between a “simple pull-out” test and measured transfer lengths. From these research programs, the precast concrete industry adopted a set of standard test procedures that were to be employed in performing “pull-out” tests. The set adopted is known as the “Moustafa Test.” Early results using the Moustafa Test indicated that the test could be used to compare the bonding characteristics of strand on a relative basis. Logan (1997) demonstrated that the Moustafa Test, at the recommended threshold value, would provide strand with bonding capability more than adequate to meet current design assumptions. In the meantime, the Post-Tensioning Institute commis- sioned a study at Queen’s University in Ontario (Hyett, Dube, and Bawden 1994). The study produced yet another bond test, the “PTI Bond Test.” The PTI Bond Test’s primary purpose was to assess the bond characteristics of 0.6 in. di- ameter strand and show the strand’s suitability for use as a rock anchor. In an appendix to ASTM A 416, the ASTM has adopted the PTI Bond Test on a provisional basis for 0.6 in. diameter strand that is to be used as rock anchors. Subsequent testing sponsored by the North American Strand Producers Association (NASP) led to the development

950 40 30 20 Z A M K B PW C J r2 = 0.92 STRESSCON=(0.597)OU+14.0 "PERFECT" TEST 10 0 0 10 20 OU DATA (kips) ST RE SS CO N DA TA (k ips ) 30 40 50 Figure 2.1. Comparison of Moustafa pull-out values from Stresscon and OU. 50 40 30 20 Z A M K B PW C J r2 = 0.88 FWC=(0.625)OU+5.25 "PERFECT" TEST 10 0 0 10 20 OU DATA (kips) FW C DA TA (k ips ) 30 40 50 Figure 2.2. Comparison of Moustafa pull-out values from FWC and OU. of a third bond test, now called the “NASP Bond Test” (Russell and Paulsgrove 1999b). In “blind trial” testing, the Moustafa Test, the PTI Bond Test, and the NASP Bond Test were performed at multiple sites. The results of the blind trial testing indicated that the NASP Bond Test provided the best repeatability. Based on these results and on as yet unpub- lished results from NASP Round III testing, the NASP recommended the use of the NASP Bond Test as the stan- dardized test to assess the bond characteristics of prestressing strands. Overall, experimental results clearly show that inherent quality differences exist in the bond of prestressing strands from various manufacturers. Accordingly, it is im- perative in a testing program to evaluate the bonding charac- teristics of the prestressing strands used. The standardization process will make possible nationwide adoption by trans- portation agencies of the experimental results on transfer length and development length of strand in concrete. Round II of the NASP tests examined the proposed stan- dardized tests for repeatability and reproducibility. The re- sults clearly indicated that the NASP Bond Test was the most reliable test of the three tests examined. Results from the Moustafa Test are shown in Figures 2.1 and 2.2. Note that in the Moustafa Test, results from a majority of strands tended to cluster near the threshold level, and a more poorly performing strand was inconsistently rated. In a similar plot, Figure 2.3 compares results from two different test series per- formed at the University of Oklahoma (OU) featuring the NASP Bond Test. Finally, Figure 2.4 compares the NASP Bond Test results at two different test sites. The repro- ducibility of test results proved to be quite remarkable and

10 25 20 15 10 Z A MK B P W C J r2 = 0.97 SERIES TWO = (0.83)SERIES ONE + 2.34 "PERFECT" TEST 5 0 0 5 10 OU SERIES ONE DATA (kips) O U SE RI ES T W O D AT A (ki ps ) 15 20 25 Figure 2.3. Comparison of NASP Bond Test results at OU in separate test series. 25 20 15 10 Z A MK B P W C J r2 = 0.97 FWC TWO = (0.658) OU ONE + 3.21 "PERFECT" TEST 5 0 0 5 10 OU SERIES ONE DATA (kips) FW C SE RI ES T W O D AT A (ki ps ) 15 20 25 Figure 2.4. Comparison of NASP Bond Test results at two different test sites. can be seen in the figures. The test has received unanimous endorsement by the NASP as its testing standard. 2.2.1.1 Effects of Strand Spacing Historically, AASHTO limited the strand clear spacing to a minimum of three times the strand diameter (3 db). In bridge codes prior to the AASHTO LRFD Bridge Design Specifica- tions, this provision was made an explicit part of the design code. It is likely that this code provision mirrored the stan- dard of placing 0.5-in. strands at 2.0 in. c/c. If this provision were extended to the larger diameter 0.6-in. strands, then the 0.6-in. strands would have to be placed at 2.4 in. c/c. Never- theless, using this strand spacing would cancel out the eco- nomic value inherent in the use of 0.6-in. strand and would also cancel out the most compelling reasons to use high- strength concrete in pretensioned girder applications. Russell (1994) showed that 0.6-in. strands must be placed at a spac- ing of about 2.0 in. c/c to enable designs to take advantage of high-strength concrete. The Auburn report (Cousins et al. 1993) was one of the more recent works dedicated to investigating the effects of strand spacing on transfer and development lengths of pre- tensioned strands. In the Auburn study, 0.5-in. pretensioned strands were fully stressed and placed at 1.75 in. c/c in some beams and 2.0 in. c/c in others. The research demonstrated that there was no substantive difference in transfer lengths measured on beams. For beams with strands spaced at 2.0 in. c/c, the measured transfer lengths averaged 44 in. For beams with strands spaced at 1.75 in. c/c, the measured transfer

11 lengths averaged 47 in. The researchers concluded that the strand spacing had no effect on the measured transfer lengths. In the same study, beams were also tested for strand devel- opment. As with the transfer length measurements, the data demonstrated that beams performed similarly regardless of whether strands were spaced at 1.75 in. or 2.0 in. c/c. The re- searchers concluded that spacing 0.5-in. strand at 1.75 in. c/c did not adversely affect transfer or development length of the strands. The researchers also concluded that the research results could be extended to the use of 0.6-in. strands at 2.0 in. c/c. From their research, Cousins et al. (1993) drew two con- clusions. First, “decreasing the strand spacing in preten- sioned, prestressed members from 2.0 inches to 1.75 inches has no significant effect on transfer length and does not result in splitting of members at transfer of prestressing force.” Second, “decreasing the strand spacing in pretensioned, prestressed members from 2.0 inches to 1.75 inches has no significant effect on development length or nominal moment capacity.” With regard to 0.6-in. strand, Cousins et al. (1993) make the following statement, “. . . for the results reported herein for specimens prestressed with 0.5 inch diameter strand, the use of 0.6 inch diameter strand at a spacing of 2.0 inches does appear reasonable.” Deatherage, Burdette, and Chew (1994) also reported on research performed to determine the effect that strand spac- ing had on transfer and development lengths. In their study, 0.5 in. diameter strand was placed in pretensioned beams with 1.75-in. and 2.0-in. spacing. Also, strands of three dif- ferent diameters (0.5 in., 0.525 in., and 9/16 in.) were placed in beams with 2.0-in. spacing. In their studies, Deatherage, Burdette, and Chew (1994) concluded that a c/c spacing of 1.75 in. should be permitted for 0.5 in. diameter strands. Also, the researchers stated that their data indicated that the bond strength of pretensioned strand was roughly proportional to its strand diameter, indicating that strand spacing did not influence the bond characteristics of strand appreciably. Accordingly, the authors recommended that the spacing requirements for 0.5-in. strand be reduced from 4.0 strand diameters to 3.5 diameters. If this principle is applied to 0.6-in. strands, the authors would effectively recommend a 2.1-in. spacing for 0.6 in. diameter strands. 2.2.1.2 Strand from Different Manufacturers Deatherage, Burdette, and Chew (1994) included 0.5 in. di- ameter strands from various manufacturers. The researchers provide strand transfer and development length test data, but provide little comment on differences between manufactur- ers. The data indicate that differences in measured transfer lengths exist among strands made by different manufactur- ers. In the Deatherage, Burdette, and Chew study (1994), the 0.5-in. strand provided by FWC (as designated in their arti- cle) had transfer lengths that varied between 18 and 36 in. Other strand manufacturers provided strand that varied between 18 in. and 21 in. In NASP Round II testing, nine different strand samples were tested. The NASP Bond Test demonstrated significant and measurable differences between strands. In the NASP Round III testing, 10 different strand samples were tested. In these tests, the differences in pull-out test results were demonstrated to correlate directly with strand transfer and development lengths. 2.2.1.3 Influence of Concrete Strength Cousins et al. (1993) also tested for transfer and develop- ment lengths in two different strength classes of concrete. The normal-strength concrete mixture resulted in concrete strengths between 6,000 and 8,000 psi. The high-strength concrete mixture resulted in concrete strengths between 10,000 and 12,000 psi. Transfer lengths measured in the high- strength concrete were, on average, 37 in.; the transfer lengths measured in the normal-strength concrete were, on average, 51 in. The higher concrete strength resulted in transfer lengths that were about 25 percent shorter. The researchers concluded that “increasing the concrete strength . . . reduces the transfer length and development length.” Two other significant research programs examined the ef- fects of concrete strength on transfer and development length. The first, undertaken by Zia and Moustafa (1977), recommended code expressions for transfer and develop- ment length that included the concrete strength parameter. Nearly 20 years later, Abrishami and Mitchell (1993) also per- formed transfer and development length tests. They also rec- ommended that concrete strength be incorporated into the code provisions. However, as noted above, the lack of data that are consistent from one research program to another has prevented the development of a consensus for code expres- sions related to transfer and development length of preten- sioned strands. 2.2.1.4 Tests of Strands Pretensioned in High-Performance Concrete In the 1990s, several research programs were undertaken by various states to design and build bridges using high- performance concrete (HPC). Most, if not all, of these proj- ects incorporated high-strength concrete as part of the HPC. In several of the projects, strand transfer length was measured, and development length tests were conducted to ensure ade- quate bonding properties from the pretensioned strands and to add to the body of knowledge regarding the transfer and de- velopment of pretensioned strands in high-strength concrete. Perhaps the first of these tests was performed in Texas by Gross and Burns (1995). In this research, two rectangular

12 beams, 42 in. deep, were fabricated. Each beam employed pre- tensioned 0.6 in. diameter strands with 2.0 in. spacing. Trans- fer lengths were measured and development length tested at each of the four ends. Concrete strengths were 7,040 psi at re- lease and 13,160 psi at the time of development length testing. From the four beam ends, an average transfer length of 14.3 in. was measured. This value is significantly less than the cur- rent transfer length provision of 60 db found in the AASHTO LRFD Bridge Design Specifications. Similarly, the development length for these 0.6-in. strands was found to be less than 78 in., which roughly corresponds to the development length given by current AASHTO provisions. The history of these beams is also interesting. They were dubbed the “Hoblitzell-Buckner” beams. Hoblitzell was employed by FHWA and was instru- mental in developing the federal programs encouraging the use of HPC. Buckner authored a report for FHWA titled, An Analysis of Transfer and Development Lengths for Preten- sioned Concrete Structures (Buckner 1994). Buckner reviewed transfer length and development length data prior to 1992/ 1993 and developed some design recommendations based on that earlier data. In his report, Buckner recommended that the design provision for transfer lengths be changed to reflect the stress in the pretensioned strand prior to release (fpi) as opposed to using the “effective prestress” after all losses, which is still found in the 318-02 Code (ACI 2002). Effectively, Buckner’s recommendation would have increased the re- quirement for transfer length by about 25 percent. More interesting was Buckner’s design equation for devel- opment length. In reviewing the data, Buckner concluded that the strain experienced by the prestressing steel at flexural strength level was an important component in the develop- ment of strand. His design equation required the design en- gineer to increase development length requirements as the steel strain at flexural strength level increased. The Hoblitzell- Buckner beams were designed, therefore, to develop ex- tremely large strains in the prestressing steel at flexural strength and test Buckner’s proposal. In the subsequent de- velopment length tests reported by Gross and Burns (1995), the strands were able to achieve their ultimate tensile capac- ity, undergo very large elongation strains, and adequately develop their tension capacities within the current AASHTO design provision. The results of these tests suggested that strand strain did not play an important role in strand devel- opment, and therefore it would not be necessary to recom- mend that the AASHTO LRFD Bridge Design Specifications should contain a development length provision based on predicted strand strain at flexural strength levels. The state of Colorado sponsored a research program specifically designed to assess the transfer length and devel- opment length of 0.6-in. strands pretensioned in HPC box beams (Cooke, Shing, and Frangopol 1998). In these beams, 0.6-in. strands were spaced at 2 in. c/c. The average measured transfer length was 23.4 in. The concrete strength at release was 7,800 psi. The box beams were also tested for development length. Concrete strength at the time of development length testing was 11,000 psi. For embedment lengths in excess of 60 in., the strands demonstrated the ability to develop adequate tension force to support the flexural capacity of the beams. Subse- quent failures were labeled as flexural failures. However, when the strand embedment length was set at 60 in. and 59 in., web shear cracking formed in the webs of the box beams, and strand anchorage failures ensued. The researchers re- ported that the development length for the strand was 60 in. Additionally, several research projects were undertaken in the 1990s in part to investigate the transfer and development length of 0.6-in. strands. Uniformly, these projects featured pretensioned 0.6 in. diameter strands and spaced at 2 in. c/c. The projects were sponsored by Texas (Barnes and Burns 2000), Virginia (Roberts-Wollmann et al. 2000; Ozyildirim and Gomez 1999), and Georgia (Khan, Dill, and Reutlinger 2002). Uniformly, these researchers concluded that 0.6 in. di- ameter strands could be deployed safely using 2-in. spacing. The state of Virginia has also supported transfer length testing of 0.6 in. diameter strains in HPC. Results reported by Ozyildirim and Gomez (1999) and Roberts-Wollmann et al. (2000) indicate that transfer lengths measured in HPC were substantially less than the transfer length predicted by the current code expressions. Barnes and Burns (2000) reported on transfer lengths that were measured on 36 AASHTO Type I beams pretensioned with 0.6-in. strands. Strand spacing was 2 in. c/c. Concrete compressive strengths at release ranged from 3,950 to 11,000 psi. Altogether, transfer lengths from 192 independent meas- urements are discussed, and the report includes data on strands that are fully bonded to the ends of the member and strands that are shielded, or debonded, at the ends of the member. The results of the Barnes and Burns study (2000) indicate a definite trend in which transfer lengths tend to de- crease in inverse proportion to the square root of the concrete strength at release. A “best fit” line reported by the authors in- cludes the square root of the concrete strength at release in the denominator. This relationship is shown in Figure 2.5. However, the data demonstrate wide variation, and the sta- tistical correlation is relatively weak. Nonetheless, it appears that concrete strength is an important factor that may affect the bond of pretensioned strand. Barnes and Burns (2000) also reported results on transfer lengths of strand from various strand manufacturers. Their results are illustrated in Figure 2.6. The data illustrated in Fig- ure 2.6 demonstrate that wide variations in measured transfer length may be the result of differences among strand manu- facturers. This finding highlights the need to establish an in- dustry standard for the “bond-ability” of prestressing strand.

13 0 10 20 30 40 50 60 AASHTO LRFD ACI/AASHTO Standard lt = 0.13 MPa-0.5 R = 0.37 70 0 Tr an sf er L en gt h (d b ) 50 100 150 (MPa0.5)fpt 200 250 300 f'ci√ fpt db f'ci√ lt = 0.22 MPa-0.5 fpt db f'ci√ 1 MPa0.5 = 0.381 ksi0.5; 1 MPa-0.5 = 2.63 ksi0.5 Figure 2.5. Data comparing transfer lengths to concrete strength at release (Barnes and Burns 2000). 0 30 60 90 120 AASHTO LRFD ACI/AASHTO Standard lt = 0.5 MPa-0.5 Manufacturer A Manufacturer B Manufacturer C Manufacturer D Unknown 150 0 Tr an sf er L en gt h (d b ) 50 100 150 (MPa0.5)fpt 200 250 300 f'ci√ db fpt f'ci√ 1 MPa0.5 = 0.381 ksi0.5; 1 MPa-0.5 = 2.63 ksi-0.5 Figure 2.6. Data highlighting differences among strand manufacturers (Barnes and Burns 2000). In addition to the research projects explicitly discussed herein, there have been other projects across the United States that have incorporated the use of 0.6 in. diameter strands and spaced at 2.0 in. c/c. Many of those projects have measured transfer lengths. One of the projects was performed by Kahn, Dill, and Reutlinger (2002). In some cases, the research reports are still in a preliminary format and use of the data is being reserved by the authors and the research sponsors. However, it is safe to say that, uniformly, these projects are employing 0.6 in. diameter strands at 2.0-in. spacing without adverse effects. 2.2.1.5 Effects of Air Entrainment There is no evidence that a systematic testing program examining the effects of air entrainment on the transfer and development of prestressing strands exists. There is a need to examine the effects of air entrainment on pretensioned bond. The research reported herein incorporates the use of air en- trainment; however, it should be noted that air entrainment is not usually specified in combination with high-strength concrete/HPC because air entrainment directly causes a decline in concrete strengths. 2.2.1.6 Water Reducers and High Range Water Reducers There is no evidence of a systematic testing program ex- amining the effects of water reducers (WRs) or high range water reducers (HRWRs) on the transfer and development of prestressing strands. Since WRs and HRWRs are used in more than 95 percent of the pretensioned prestressing plants throughout North America, this is an important variable that warrants investigation.

14 2.2.2 Development and Splice Length for Mild Reinforcement To identify needed experimental research, the literature re- view focused on the analysis of test results from bond tests on development and splice length in tension of coated and un- coated bars and development length of coated and uncoated bars terminated with standard hooks in tension. Based on the reported bond performance of individual and bundled bars in compression, it was determined that no additional exper- imental work was required in this area. Compression devel- opment lengths are considerably shorter than tension development lengths because there are no transverse cracks in compression zones; the harmful effect of such cracks in ini- tiating splitting is absent. However, the major difference be- tween tension and compression development and splice lengths is the ability of the bars in compression to transfer load to the concrete directly by bearing. In tests conducted by Pfister and Mattock (1963), bearing stresses equal to five times the cylinder strength of the concrete were attained at the square-cut ends of bars in compression splices. Addi- tional experimental work conducted at the Otto-Graf- Institute of the University of Stuttgart by Leonhardt and Teichen (1972) conclusively showed the following: • End bearing is responsible for the majority of splice failures in compression irrespective of the splice length tested. The splice lengths varied between 9 and 38 bar diameters. • The bearing capacity of the concrete at the bearing ends of the bars was increased by the presence of confining rein- forcement. Under such conditions, concrete bearing stresses of 17 ksi were measured (for concrete with a uni- axial compressive strength around 4 ksi). • An increase in the thickness of the concrete cover over the compression splice resulted only in very minor improve- ments in bond performance. • Under long-term loading, the bearing pressure under the ends of the compression bars diminishes because of creep; hence, the splice performance improves. The available information on the anchorage in tension of welded wire reinforcement indicated that a significant exper- imental effort was not required as part of NCHRP Project 12-60 (Furlong, Fenves, and Kasl 1991; Griezic, Cook, and Mitchell 1994; and Guimaraes, Kreger, and Jirsa 1992). In the case of plain wire fabric, the development in tension depends on the mechanical anchorage from at least two cross wires. Deformed welded wire reinforcement derives anchorage from bond stresses along the deformed wires and from mechanical anchorage from the cross wires. Current code ex- pressions for development length in tension of deformed welded wire reinforcement assume that at least one cross wire is present in the development length. Tests have also shown that the development length of deformed welded wire rein- forcement is not affected by epoxy coating, and thus the epoxy coating factor in the current ACI Code is 1.0 for epoxy- coated deformed wire fabric. In recent years, welded wire fabric (WWF) has been used widely as shear reinforcement in thin-webbed girders because of the ease of construction over the use of conventional stirrups. Research conducted to date, with concrete compressive strengths up to 12 ksi, indicates that this reinforcement can be used effectively to resist shear (Mansur, Lee, and Lee 1987; Xuan, Rizkalla, and Maruyama 1988; Pincheira, Rizkalla, and Attiogbe 1989; and Zhongguo, Tadros, and Baishya 2000). It was shown that two cross wires welded at a spacing of 2 in. at the open ends (top and bottom) of WWF cages provide satisfactory anchorage. Such anchor- age was found to be more effective for deformed WWF than smooth WWF. The increase in concrete compressive strength has been shown to further improve the anchorage of this reinforcement. 2.2.2.1 Databases There are two databases. One consists of 71 tension devel- opment and splice tests of specimens with top cast uncoated reinforcing bars, 493 specimens with bottom cast uncoated reinforcing bars, 27 specimens with top cast epoxy-coated bars, and 48 specimens with bottom cast epoxy-coated bars, for a total of 639 specimens. The other database consists of 33 specimens with uncoated bars terminated with standard hooks and 13 specimens with epoxy-coated bars, for a total of 46 specimens. The provisions for development length of reinforcement in Section 5 of the AASHTO LRFD Bridge Design Specifications are based on the provisions of ACI 318-89 (ACI 1989). The 1989 provisions in the ACI Code were extensively modified in the 1995 version of the ACI Code (ACI 1995) with a view to formulating a more “user-friendly” format while main- taining the same general agreement with professional judg- ment and research results. Tests conducted by Azizinamini et al. (1993, 1999a) have indicated that in the case of high- strength concrete, some minimum amount of transverse re- inforcement is needed to ensure adequate ductility from the splice at failure. A proposed modification to ACI 318-99 (ACI 1999), based on these tests, called for the determination of a basic straight development length for bars in tension without including the presence of transverse reinforcement, together with a minimum area of transverse steel in the form of stir- rups, Asp, crossing potential splitting planes. In these studies, over 70 specimens with concrete compressive strengths rang- ing between 5 ksi and 16 ksi were tested (Azizinamini et al. 1993, 1999a).

15 0.00 0.20 0.40 0.60 0.80 1.00 1.20 1.40 1.60 0 2000 4000 6000 8000 10000 12000 14000 16000 18000 f'c (psi) U t es t (ks i) Bottom Cast Uncoated Figure 2.7. Bond stress at failure (utest) versus the concrete compressive strength ( ) of bottom cast uncoated specimens. ′fc Although the modification proposed in another paper by Azizinamini and colleagues (1999b) was not adopted in the 2002 version of the 318 Code (ACI 2002), it was deemed an improvement over the current AASHTO LRFD provisions. Therefore, the 2005 318 Code (ACI 2005) provisions were used in NCHRP Project 12-60 as the basis for further exten- sion of the AASHTO provisions to higher strength concrete. The experimental work conducted in the mild steel phase of NCHRP Project 12-60 was focused on filling the gaps identi- fied in order to extend the applicability of the present AASHTO LRFD Bridge Design Specifications to normal- weight concrete with compressive strengths up to 15 ksi. The 639-specimen database is shown in Figures 2.7 through 2.10 by plotting the bond strength, utest, versus the concrete compressive strength, f ′c. The bond strength is defined as (2.1) In Equation 2.1, Ab is the area of bar being developed or spliced, fsu is the stress in the bar estimated at failure using moment-curvature type analysis and compatibility of defor- mations, db is the diameter of the bar, and ls is anchorage/ splice length. As can be seen from Figures 2.7 and 2.8, there is a lack of data for development and splice lengths of un- coated bars in tension above 16 ksi. Figures 2.9 and 2.10 show that there are limited data for epoxy-coated bars in tension u A f d l b su b s test = π above 8 ksi. In order to assess whether the limit on f ′c can be removed by examining the existing data for development and splice length of uncoated and epoxy-coated bars in tension, the ratio of test to calculated bond strength is plotted versus the concrete compressive strength (f ′c) evaluated throughout the range of concrete cylinder strengths. The bond strength ratio is determined in terms of bar stresses at failure versus calculated bar stress, using Equations 2.2 through 2.5: (2.2) (2.3) (2.4) To limit the probability of a pull-out failure, 318 Code (ACI 2005) requires that (2.5) The additional parameters in the equations are the follow- ing: fs is the stress in the reinforcing bar; cmin is the smaller of minimum cover or one-half of the clear spacing between bars; Atr represents the area of each stirrup or tie crossing the potential plane of splitting adjacent to the reinforcement c K d tr b + ≤ 2 5. K A f s n tr tr yt = * *1500 * c c db= +( )min . *0 5 l d f f c K d s b s c tr b = ′ +⎛⎝⎜ ⎞⎠⎟ 3 40

16 0.00 0.20 0.40 0.60 0.80 1.00 1.20 1.40 1.60 0 2000 4000 6000 8000 10000 12000 14000 16000 18000 f'c (psi) u te st (ks i) Top Cast Uncoated Figure 2.8. Bond stress at failure (utest) versus the concrete compressive strength ( ) of top cast uncoated specimens. ′fc 0.00 0.10 0.20 0.30 0.40 0.50 0.60 0.70 0.80 0.90 1.00 0 2000 4000 6000 8000 10000 12000 f'c (psi) u te st (p si) Epoxy-Coated Bottom Bars Figure 2.9. Bond stress at failure (utest) versus the concrete compressive strength ( ) of bottom cast epoxy- coated specimens. ′fc

17 0.00 0.10 0.20 0.30 0.40 0.50 0.60 0.70 0.80 0 2000 4000 6000 8000 10000 12000 14000 f'c (psi) u te st (ks i) Epoxy-Coated Top Cast Figure 2.10. Bond stress at failure (utest) versus the concrete compressive strength ( ) of top cast epoxy-coated specimens. ′fc being developed, spliced, or anchored; fyt is the yield strength of the stirrup reinforcement; s is the spacing of stirrups; and n is the number of bars being developed or spliced. The re- sults of the evaluation indicated that the average of the ratio for bars not confined by stirrups is 1.23 with a standard devi- ation of 0.28 for all f ′c values, and 1.23 with a standard devia- tion of 0.23 for concrete compressive strengths below 10 ksi. In the case of bars confined by stirrups, the average is 1.23, and the standard deviation is 0.3 for all f ′c values. For f ′c values below 10 ksi, the average is 1.24 and the standard deviation is 0.30. In members with confined bars, the stirrups are as- sumed to be uniformly spaced throughout the splice/devel- opment length. The value for the members in the database, calculated by the ACI provisions, gives approximately the same scatter throughout the range of concrete compressive strengths up to a maximum of 16 ksi for members with and without stirrups. This conclusion supports the extension of these provisions to higher concrete compressive strengths with a few verification tests of uncoated bars at the upper limit, mainly to establish the role of the minimum amount of transverse reinforcement on the mode of failure of splices in tension recommended in the Azizinamini et al. studies (1993, 1999a). On the other hand, it is recognized that there is a paucity of data on the performance of epoxy-coated bars in concretes with compressive strengths above 10 ksi. Therefore, a more intense verification testing effort was carried out in this study to close this gap. Tests have shown that the bar force is transferred rapidly into the concrete, and the portion following a hook is gener- ally ineffective and can potentially be limited by the tensile strength of the concrete. Marques and Jirsa (1975) reported on the results of 22 tests conducted using two #7 or two #11 uncoated bars. Standard 90- or 180-deg hooks conforming to the 318 Code were used (ACI 2005). The concrete compres- sive strength was around 5 ksi. The specimens simulated exterior beam column joints. Hamad, Jirsa, and D’Abreu de Paulo (1993) reported on the results of 24 tests to evaluate the anchorage performance of epoxy-coated hooked bars. Based on these results, a 20-percent increase on the basic develop- ment length was recommended for epoxy-coated hooked bars. It was shown that the relative anchorage strength of un- coated and epoxy-coated hooked bars was independent of bar size, concrete strength, side concrete cover, or hook geome- try. The maximum concrete strength of the specimens was 7 ksi. These tests serve as the basis of the 318 Code anchorage provisions for bars anchored by means of standard hooks (ACI 2005). The specimen and the test setup used in NCHRP Project 12-60 was similar to the one used in the Marques and Jirsa (1975) and Hamad, Jirsa, and D’Abreu de Paulo (1993) studies. However, only 90-deg hooks were evaluated, since Hamad, Jirsa, and D’Abreu de Paulo found little difference in the performance of 90- and 180-deg hooks. It should be noted that sections of the AASHTO LRFD Bridge Design Specifications dealing with the anchorage of bars in tension

18 Cc Cs = min{C1, C2/2} C1 C2 (b) (c)(a) Figure 2.12. Anchorage failure modes: (a) vertical splitting, (b) splitting in the horizontal plane of the bars, and (c) pull-out without splitting (ACI 408 2003). bearing and friction forces on bar adhesion and friction forces along the surface of the bar Figure 2.11. Mechanisms of force transfer between concrete and reinforcement- deformed bars. terminated with a standard hook are similar to those in the 2005 version of the ACI Code. 2.3 Identification of Issues and Needs The work described in the previous section was used to as- semble a comprehensive list of issues pertaining to transfer length, development length, and splice length of strand/ reinforcement to normal-weight concrete with compressive strengths in excess of 10 ksi and up to 15 ksi. In this section, a discussion of the main issues related to bond performance of reinforcement is presented, and gaps found in the existing data- base are addressed. The experimental program described in Chapter 3 of this report was directed at addressing the identi- fied needs in order to extend the AASHTO LFRD Bridge Design Specifications to allow greater use of high-strength concrete. In reinforced and prestressed concrete structures, suffi- cient transfer of forces between concrete and reinforcement is required for a satisfactory design. The transfer of forces occurs through a combination of chemical surface adhesion, friction, and bearing of bar deformations against the sur- rounding concrete. Initially, the transfer of forces occurs mainly by chemical adhesion; after initial slip, most of the force is transferred by bearing and friction. In the case of plain bars or wires, slip-induced friction—resulting from shape and surface roughness—plays an important role in the force transfer. In the case of deformed reinforcement, as slip in- creases, bearing of the ribs against the surrounding concrete becomes the principal mechanism of force transfer between concrete and steel. The forces on the bar surface are balanced by compressive and shearing stresses in the concrete (see Fig- ure 2.11). The concrete stresses result in tensile stresses that, if high enough, can lead to cracking in planes both parallel and perpendicular to the reinforcement, as shown in Figure 2.12. These transverse cracks can lead to splitting failure. If the concrete cover, bar spacing, or amount of transverse reinforcement is sufficient to prevent or delay the splitting fail- ure, then failure can occur along a surface surrounding the perimeter of the bar, resulting in a pull-out type failure. Tests have shown that these two types of failures can take place at stresses close to the tensile strength of the reinforcement. Pull- out failures occur in cases of high confinement and low bonded lengths. However, splitting failures are more common in structural applications. For this reason, it is recommended that experimental data considered for development of design equations should have a minimum embedment length. Another important observation is that transverse reinforce- ment has been observed to rarely yield during splitting failures (Maeda, Otani, and Aoyama 1991; Sakaruda, Morohashi, and Tanaka 1993; and Azizinamini et al. 1999a). Therefore, it is important to limit in design provisions the level of confine- ment provided by transverse reinforcement. The many factors affecting bond performance are presented in two main cate- gories: member properties and material properties. Some of the factors are common to both strand and mild reinforce- ment while others are unique to one or the other. Initially, in the testing of prestressing strand for bond to con- crete, the simple pull-out tests were criticized because they did not include the wedging action, or Hoyer’s effect, associated with pretensioned strands in real beams. However, subsequent testing with both the Moustafa Test and the NASP Bond Test have demonstrated that a direct correlation exists between re- sults from these simple pull-out tests and strand performance in pretensioned beams. Therefore, in this testing program the NASP Bond Test was employed as an assessment tool to quan- tify the “bond-ability” of prestressing strands that will be employed. Testing sponsored by the NASP has demonstrated that the NASP Bond Test has superior repeatability and repro- ducibility when compared with the Moustafa Test. 2.3.1 Member Properties 2.3.1.1 Transfer Length of Prestressing Strand In the specific case where prestressing strands are bonded to concrete, bond stresses are derived through a combination of adhesion, friction, and mechanical interlocking (Hanson

19 Figure 2.13. Variation of steel and bond forces in a reinforced concrete member subjected to pure bending: (a) cracked con- crete region, (b) bond stresses acting on a reinforcing bar, (c) variation of tensile force in steel, and (d) variation of bond force along the bar (Nilson 1997). and Kaar 1959). It has been widely believed that a wedging effect, called Hoyer’s Effect, unique to pretensioned strands, creates significant bond stresses in the transfer zone where the effective prestressing force is transferred from the preten- sioned strand to the concrete. In those same regions, slip occurs between strand and concrete due to the difference in strain condition. Research has indicated that the strand end slip can be used as a quality control measure for the bond of prestressing strands (Rose and Russell 1997). Furthermore, the relative slip between strand and concrete virtually ensures that adhesion plays little or no role in the transfer of pre- stressing forces to concrete (Russell and Burns 1996). 2.3.1.2 Development and Splice Length Bond forces are not uniformly distributed over the length of anchorage (see Figure 2.13). Thus, bond failures are incremen- tal, initiating in the region of highest bond force per unit of length. In the case of anchored reinforcement by means of straight embedment, longitudinal splitting will initiate at either a free surface or at a flexural crack location. In the case of spliced bars, splitting will start at the ends of the splice and move toward the center. The mode of failure explains the fact that the non- loaded end of a developed/spliced bar is less effective than the loaded end in transferring forces between concrete and rein- forcement. It can be concluded that there is a non-proportional relationship between development/splice length and bond strength. Thus, even though bond strengths have been meas- ured for very short embedment lengths, it is not appropriate to linearly extrapolate such findings to code development lengths. This observation suggests the need for testing at appropriate scale for development of design provisions. In beams tested for strand development, it is equally appar- ent that cracking causes the mobilization of the strand relative to concrete. Commonly, bond stresses that develop strand tension from the transfer zone to the point where flexural capacity is required are called “flexural bond stresses.” Flex- ural bond results primarily from a combination of mechani- cal interlocking and friction. The mechanical interlocking bond stresses are derived by the helical windings of the 7-wire prestressing strand, which act similarly to the mechanical de- formations found on rolled, mild reinforcement. Development length testing of pretensioned beams indi- cates that splitting occurs less frequently than in convention- ally reinforced beams (although splitting cracks have been observed in pretensioned bond failures). Issues for strand development are more related to the cracking patterns that occur as the pretensioned beams approach their ultimate strength. In testing on beams with debonded strands, it is clear that cracks that propagate through or near the transfer zone of pretensioned strands cause strands to slip. In many of those tests, cracking in the transfer zones of pretensioned strands caused bond failure of pretensioned strands (Russell, Burns, and ZumBrunnen 1994; Russell and Burns 1994). Additionally, in pretensioned strands with fully bonded beams, it is important to note that sections with narrow webs, specifically I-shaped beams, have failed in bond in concert with web shear cracking that occurs near or through the transfer zones of pretensioned strands (Jacob 1998; Kaufman and Ramirez 1988; Russell and Burns 1993). In contrast, tests on rectangular prestressed beams will not produce web shear cracks, so the behavior of rectangular cross sections can be significantly different than cross sections with narrow webs. For that reason, the testing program includes testing of both rectangular and I-shaped sections. 2.3.1.3 Transverse Reinforcement Orangun, Jirsa, and Breen (1977) indicated that transverse reinforcement confines the concrete around anchored bars

20 and limits the progression of splitting cracks. An additional beneficial effect of transverse reinforcement is that increases in transverse reinforcement lead eventually to pull-out failures rather than splitting-type failures. However, the Orangun, Jirsa, and Breen study also noted that transverse reinforce- ment in excess of the amount required to cause the change in mode of failure is not as effective and eventually leads to no further increase in bond strength. These observations and the observations by Maeda, Otani, and Aoyama 1991; Sakaruda, Morohashi, and Tanaka 1993; and Azizinamini et al. 1995 that the transverse reinforcement confining the anchored bar sel- dom yields in splitting failure indicates the need for an upper limit on the improvement in bond strength provided by the presence of transverse reinforcement. 2.3.1.4 Casting Position It has been observed by various researchers that top cast bars have lower bond strengths than bottom cast bars. Clark (1946), using pull-out type specimens cast in a horizontal position, noted that in the top position, bars were two-thirds as effective in bond as in the bottom position. The depth of the concrete under the bar in the top position was 15 in., and the depth of the concrete under the bar in the bottom posi- tion was 2 in. The concrete slump was 4.25 in., and the com- pressive strength averaged 5.6 ksi. Ferguson and Thompson (1965) noted that with 12 in. of concrete below the bar, the strength dropped from 3 to 13 percent as the slump was in- creased. They noted that for the beam depths tested, from 13 to 22 in., the 1.4 factor used in the specifications was conser- vative. This observation is currently recognized in the 318 Code where a 1.3 factor is used to increase the development length or splice of bars cast horizontally with more than 12 in. of fresh concrete cast in the member below the bar (ACI 2005). A 1.4 factor is currently prescribed in the AASHTO LRFD Bridge Design Specifications to cover this case, and it is thus conservative if the effects of cover and transverse rein- forcement are included in the specifications. Additional research (Jirsa and Breen 1981) indicates that the concrete slump plays an important role in determining the effects of casting position, and this is most significant when very large depths of concrete are cast below the bars or splices. The 1981 study by Jirsa and Breen further indicated that the so-called top bar factor should vary with the depth of concrete cast below the reinforcement and recommended a maximum factor of 1.3 for slumps of less than 4 in. For slumps between 4 and 6 in., a maximum factor of 1.35 is rec- ommended for depths below 24 in., and a maximum factor of up to 1.6 is recommended for depths greater than 48 in. For slumps greater than 6 in., a factor of 1.8 for depths below 24 in. is recommended, and a factor of 2.2 for depths below 48 in. is recommended. It is further stated that the basic bond strength of vertical bars seems to be reduced only by 25 per- cent with respect to the bond strength of horizontal bars. A single factor of 1.3 is recommended for all vertical bars where the center of the splice or the development length has more than 24 in. of concrete cast below. 2.3.1.5 Concrete Cover and Spacing of Reinforcement As shown in Figure 2.12, splitting failure is expected to con- trol in the majority of structural applications. In this type of fail- ure, the actual location of the splitting cracks in the case of bot- tom cast reinforcement depends on the relative values of the concrete bottom cover, concrete side cover, and one-half of the clear spacing between bars. If the bottom cover is less than the side cover and one-half the spacing between bars, splitting occurs through the cover to the bottom free surface. If either the side cover or one-half the bar spacing is smaller than the bot- tom cover, then splitting of the concrete occurs either through the side cover or between the reinforcement. This observation supports the need to modify the current AASHTO LRFD Bridge Design Specifications for bond and development length of mild reinforcement to incorporate the effects of cover, bar spacing, and transverse reinforcement. 2.3.2 Material Properties 2.3.2.1 Reinforcement Properties For a given bonded length required to achieve a given steel stress level, reinforcement of different areas will achieve dif- ferent levels of force at the onset of splitting failure, with the larger area reinforcement achieving higher forces. Therefore larger area reinforcement will require longer development/ splice length than smaller area reinforcement for the same degree of confinement. The size of the reinforcement being developed also plays a role in the contribution of the confin- ing reinforcement for the case of deformed bars. As large bars slip, higher strains are mobilized in the transverse reinforce- ment, thus the beneficial effect of transverse reinforcement on the bond strength of deformed bars increases as the area of the bar increases. It is now customary to relate bond performance to bar geometry by means of the relative rib area factor, Rr, defined as: (2.6) Typical bars currently used in the United States have relative rib area factors ranging between 0.057 and 0.087 (Choi et al. 1990). Darwin and Graham (1993a, 1993b) concluded that the bond strength is independent of deformation pattern if the bar Rr = projected rib area normal to bar axis (nom. bar perimeter) (center to center rib spacing)× − −

21 is under small cover conditions and there is no transverse rein- forcement. Darwin and Graham observed that under large cover or with transverse reinforcement, bond strength in- creased with an increase in relative rib area. They also found that deformations parallel to the splitting cracks were more effective. The bond strength of epoxy-coated bars has been found to increase under all conditions of confinement as the relative rib area is increased. Zuo and Darwin (1998) recommended that for epoxy-coated bars with relative rib areas greater than or equal to 0.1 and concrete with compressive strength below 10 ksi, development and splice length should be increased by 20 percent instead of the 50-percent increase for cover less than 3db or clear spacing less than 6db. For concrete strengths greater than 10 ksi, a 50-percent increase appeared warranted regardless of the value of Rr. The surface condition is important from the standpoint of bond strength because it affects adhesion, friction, and bear- ing in the transfer of forces between steel and surrounding concrete. Items such as cleanliness, rust, and coatings affect the surface condition of the reinforcement. Specifications require that the reinforcement be free of mud and other substances capable of reducing bond strength. It is well established that the presence of epoxy coatings reduces the bond strength of reinforcement (Mathey and Clifton 1976; Johnston and Zia 1982; Treece and Jirsa 1989; Choi et al. 1990, 1991; Cleary and Ramirez 1993). 2.3.2.2 Concrete Properties Compressive strength and lightweight aggregate are ac- knowledged in codes and specifications as influencing bond strength. In addition, tensile strength and fracture energy, mineral admixtures, and consolidation and vibration are also factors affecting bond strength of reinforcement. Azizinamini et al. (1993, 1999a) noted that for higher strength concretes, the higher bearing capacity prevents crushing of the concrete in front of the ribs, thus reducing the local slip. These researchers further noted that the reduced slip also limited the number of ribs participating in the load transfer between concrete and reinforcement. The reduced participation of the ribs increases the local tension stresses and further leads to a non-uniform distribution of bond force. Although traditionally has been used to reflect the concrete compressive strength in bond calculations, Zuo and Darwin (1998, 2000) have postulated that f c′1/4 for members without stirrups and f c′3/4 for members with stirrups better re- flect the effect of concrete strength on bond. These findings indicate that if bond strengths are normalized with respect to f c′1/2, the effect of concrete strength on the bond strength is se- verely overestimated. High-strength concrete has been shown to improve anchorage of prestressing strand, thus reducing the required transfer length and development length. ′fc In unconfined bar bond tests, the use of basalt aggregates has been shown to increase bond strength by almost 13 percent over the bond strength of concretes with weaker aggregate such as limestone (Zuo and Darwin 1998, 2000). Tests on bars con- fined by transverse reinforcement (Darwin et al. 1996; Zuo and Darwin 1998) also indicate an increase in bond strength in the presence of stronger aggregates showing a significant effect on the contribution from the transverse reinforcement. Lower strength aggregates, on the other hand, have a detrimental effect on the bond strength. Reports by ACI Committee 408 (1966, 1970) have emphasized the paucity of experimental data on the bond strength of reinforced con- crete elements made with lightweight aggregate concrete. The AASHTO LRFD Bridge Design Specifications includes a fac- tor of 1.3 for development length to reflect the lower ten- sile strength of lightweight aggregate concrete and allows that factor to be taken as 0.22 if the average splitting strength, fct, of the lightweight aggregate concrete is specified. For lightweight sand, where fct is not specified, a factor of 1.2 is specified. Although design provisions, in general, require longer development lengths for lightweight aggregate con- crete, test results from previous research are contradictory, in part, because of the different characteristics associated with the particular type of aggregate and mix design. The use of lightweight aggregate concrete is outside the scope of NCHRP Project 12-60. It has been widely observed that as the concrete compres- sive strength increases, the bond strength of the same con- crete also increases—albeit at a slower rate—leading to po- tentially more brittle failures (Azizinamini et al. 1993, 1999a). On the other hand, the tensile strength of the concrete is not the only factor controlling bond strength, as it has been noted by Zuo and Darwin (1998, 2000) for deformed bars. The Zuo and Darwin studies recommended the use of f c′1/4 instead of the traditional f c′1/2 to represent the effect of concrete com- pressive strength on bond strength for unconfined bars. They also noted that the presence of confinement influenced the power of the compressive strength and recommended the use of f c′3/4 as a good representation of the influence of compres- sive strength on bond strength. Most of the work related to bond has focused on the effect of silica fume. The studies have shown increases of less than 10 percent on bond strength in the presence of the mineral admixture (DeVries, Moehle, and Hester 1991; Hamad and Itani 1998). 2.4 Issues Related to Testing Protocols A review of testing protocols for determining bond charac- teristics was presented. From our review of available research, we recommended that the NASP Bond Test be employed ′ ≥f fc ct/ .1 0

22 throughout the experimental program to quantify the bond characteristics of individual strand samples. The testing program includes “round robin” testing at both Purdue Uni- versity and Oklahoma State University (OSU) to validate the repeatability of the test procedure. The NASP Bond Test procedure has been refined through this research and is now recommended for adoption into the AASHTO LRFD Bridge Code as the Standard Test Method for the Bond of Prestress- ing Strands. The testing protocols for splice/development lengths and bars terminated with standard hooks are also pre- sented. The beam-splice test for splice/development length and the Marques and Jirsa (1975) and Hamad, Jirsa, and D’Abreu de Paulo (1993) exterior beam column joint setup for bars in tension anchored by means of standard hooks are recommended for use in this study. 2.4.1 Testing Protocols for Prestressing Strand Since 1994, three new test procedures or protocols have been developed for assessing the bonding characteristics of prestressing strand: the Moustafa Bond Test, the Post- Tensioning Institute (PTI) Bond Test, and the NASP Bond Test. Testing has demonstrated that the NASP Bond Test de- livers the greatest degree of repeatability and reproducibility of the three tests. Therefore, the testing program for NCHRP Project 12-60 employed the NASP Bond Test as the standard test to assess the relative “bond-ability” of prestressing strands. Previous experience with research on strand bond demonstrates the importance of quantifying the strand bond- ing properties prior to or concurrent with testing programs for transfer and development length of strands. 2.4.1.1 Prestressing Strand up to 0.6 in. in Diameter Engineers and contractors concerned with the bond of pre- stressing strand used for rock anchors developed the PTI Bond Test. In conformance with standard practice for rock anchors, the test protocol indicates explicitly that testing should be con- ducted on 0.6-in. strand. The test protocol has been modified to accept 0.5-in. strand, but the acceptance value has not been adjusted or evaluated using 0.5-in. strand. Both the Moustafa Test and the NASP Bond Test were developed using 0.5-in. strand. In the experimental program, the NASP Bond Test was performed using 0.6-in. strand. The testing demonstrated that the Standard Test Method for the Bond of Prestressing Strands is suitable for 0.6-in. strands as well as 0.5-in. strands. 2.4.1.2 Influence of Concrete Strength As noted in the literature review, concrete strength has long been described as an important variable affecting the transfer and development of prestressing strands. Through the years, several researchers have included concrete strength as a variable in transfer and development length equations. However, the lack of consistency in the strand products themselves has worked against developing a consensus re- garding the effect of concrete strength. By quantifying strand bond characteristics through the NASP Bond Test, this re- search has been able to assess the effects of concrete strength on transfer and development of pretensioned strands. Recommendations to include concrete strength in the design equations have been made. 2.4.1.3 Influence of Water-Reducing Admixtures There is a lack of data available to assess what effect, if any, water-reducing admixtures have on the bond of pretensioned strands. This contrasts directly with the fact that more than 99 percent of the prestressing plants in North America use HRWRs (the source for this information is an informal, unpublished committee report on a survey of precast/ prestressing plants done by the Prestressing Steel Committee of PCI circa 1998). For historical perspective, it is noted that the bulk of development regarding the Moustafa Test em- ployed concrete that did not contain HRWRs. A majority of the Moustafa testing has been performed at Stresscon Corporation in Colorado Springs, where HRWRs are not commonly employed. Yet, others that have participated in Moustafa Testing have employed HRWRs as part of the stan- dard casting procedures used in the local prestressing plants. Variations that result from the use of HRWR have not been measured or quantified. These data were compiled informally through the work of the Prestressing Steel Committee of the Precast/Prestressed Concrete Institute and are not available for publication. 2.4.1.4 Influence of Air Entrainment There is a lack of data available to assess what effect, if any, air entrainment has on the bond of pretensioned strands. The experiences of the states are mixed with regard to whether air entrainment is required in pretensioned beams. The NASP Bond Test was employed to examine what effects, if any, air entrainment has on bond. Results indicate that concrete strength is more important to bond strength than air entrainment. 2.4.1.5 Top Bar Effects There is a small database in existence available to examine the “top bar effect” on transfer and development of pre- stressing strands. This information is primarily available from testing programs on prestressed concrete piling. The top bar

23 (a) (b) (e) (g) (h) (i) (j) (k) (l) (m) (d) (c) (f) P = applied load. Figure 2.14. Testing methods to evaluate bond strength. effect is expected to be assessed during casting of the scale model and full-sized specimens by including pretensioned strands in the top half of the cross section. 2.4.2 Testing Protocols for Mild Reinforcement A review of testing protocols was conducted to determine the appropriate testing protocol(s) for addressing gaps in the experimental data. It is well established that testing protocols to evaluate development and splice length requirements for deformed bars and wire in tension must be of an appropriate scale, containing more than one bar or wire; testing protocols should also show due regard for a realistic transfer of force between concrete and steel reinforcement, as well as cover and bar spacing effects. The more commonly used testing configurations are shown in Figure 2.14. Although they are economically appealing, pull-out tests used by earlier researchers to evaluate bond performance of various reinforcing bars embedded in concrete of different strengths (Figures 2.14[a] through [e]) present the problem of intro- ducing transverse compression, a compression not typical of situations encountered in structures. Transverse compression has a beneficial effect on bond strength and yields an overly optimistic assessment of the actual performance of struc- tures. For this reason, various testing schemes have been pro- posed to eliminate transverse compression (see Figure 2.14[f] and [g]). In the case of semi-beam specimens, such as those shown in Figure 2.14(f), it is critical to properly account for the increase in the length over which splitting resistance tends to be mobilized due to the confining pressure at the end of the bar (if the bar end is not shielded). ACI Committee 408 (1964) prepared a detailed guide for the determination of bond strength in beam specimens. The more popular varia- tion, the so-called beam splice test with the splice located in the constant moment region (the most critical condition is one where both bars in the splice are subjected to high stresses), can be seen in Figure 2.14(i). Splice tests have been realistic simulations of real conditions in structures, but development length tests have been con- ducted largely using pull-out tests in which splitting failures are purposely avoided. As a result, the bond stresses developed along splices are low compared with the bond along a bar in a pull-out test. This difference in test methods is responsible for

24 Figure 2.15. Exterior beam-to-column joint setup to evaluate bond performance of bars developed using standard hooks. large differences in code-required anchorage lengths for splices and development of single bars. Pull-out failures occur in cases of high confinement and short bonded lengths. In most structural applications, however, splitting failures tend to control. On this basis, the data developed for extending the current AASHTO LRFD specification for splice/development length had a minimum bonded length to bar diameter ratio of 15 in order to avoid unrealistically high values of bond strength. A similar concept of minimum embedment length should be included in any proposed specifications. Splice specimens such as those shown in Figure 2.14(i) are deemed to represent larger-scale specimens designed to directly measure development and splice strength in full-scale mem- bers. Because of the relative ease of fabrication and the realistic state of stress achieved during testing, splice specimens were used in the development of experimental data on development/ splice length of mild reinforcement in this research. Review of experimental data on anchorage of bars termi- nated using standard hooks indicates the need for additional testing to extend the current AASHTO LRFD specifications to concrete strengths up to 15 ksi (see Section 2.2.2). Tests have shown that the bar force is transferred rapidly into the con- crete, and the portion following a hook is generally ineffective and can potentially be limited by the tensile strength of the concrete. Further study of failures of hooked bars indicates that splitting of the concrete cover is the primary cause of failure and that splitting originates at the inside of the hook, where the local stress concentrations are higher. Thus, it has been deter- mined that hook development is a direct function of bar di- ameter, db, which governs the magnitude of compressive stresses on the inside of the hook. The experimental work sup- porting the current requirements for development of standard hooks in tension was conducted using the test setup shown in Figure 2.15. In NCHRP Project 12-60, a similar specimen and test setup was used in the evaluation of uncoated and epoxy- coated bars terminated with standard hooks in tension to normal-weight concrete with compressive strength up to 15 ksi. A useful test protocol to help understanding the bond strength of mild reinforcement in concrete members must define a minimum level of information to be provided. The recommended level of information is described in the following subsections. 2.4.2.1 Concrete Properties The following information on concrete properties should be provided: • the source of the concrete. • the mix proportions, including identification of the components: – cement type; – mineral admixtures; – chemical admixtures, including specific gravity and percent solids; and – fine and coarse aggregates and their properties (e.g., specific gravity [SSD] and absorption). • the concrete compressive strength, as obtained from a standard concrete cylinder (which should be cured side- by-side with, and in the same manner as, the bond/splice specimens), and including: – size of the compressive strength specimens, – type and thickness of the cylinder caps used on the spec- imens, and – age of the specimen at testing. • the concrete flexural strength, including: – size of the flexural strength specimens, – age of specimen at testing, and – flexural test method used. 2.4.2.2 Reinforcement Properties The properties of the reinforcing steel are required for basic identification and, in most cases, are needed to fully characterize the steel used in the tests. The following infor- mation should be provided for each heat or production run of reinforcing steel: the standard (ASTM) under which the bars were manufactured, the nominal diameter, bar designa- tion, yield strength, tensile strength, proof strength (if appli- cable), elongation at failure, weight (mass) per unit length,

25 rib spacing and rib height (according to the standard under which the bar was manufactured as well as the average value), relative rib area, rib angle (the included angle between the rib and the bar axis), rib-face angle, and type of coating and coat- ing thickness (if applicable). 2.4.2.3 Specimen Characteristics The following specimen characteristics should be recorded: exterior dimensions, the location of structural/tensile rein- forcement (including the effective depth), bottom/top cover, side cover, the clear spacing between bars, and the length of the specimen. These additional specimen characteristics should be recorded: the length of the developed/spliced bars, the number of developed/spliced bars, the number and average spacing of stirrups/ties used as transverse reinforcement in the region of the developed/spliced bars, the tensile strength of this rein- forcement, and the load (tensile/bending) on the specimen at the time of failure (including the specimen self-weight and the weight of the test system). Specimen dimensions should be measured after casting. It is also recommended that the cover be measured after casting and/or testing. 2.4.2.4 Test Information The following basic information should be recorded for each test: a description of the test system; the weight of the loading system; the rate of loading; the presence or absence of strain gages on the developed/spliced reinforcing bars; loca- tion, type, and number of displacement transducers; the full load-deflection curves for the tests; and the number and location of load measurement devices. More information on the local behavior of the splice or development length can be obtained by placing strain gages on the bar itself. The strain gage instrumentation provides information on the changes in bar force along its length. In order to avoid excessive distur- bances due to the presence of the gages, they are installed by splitting the bar in half, forming a channel along the center- line. Strain gages and wires are then placed in the channel, and the bar is welded together. An acceptable alternative is to place the wires and gages in grooves cut along the longitudi- nal ribs of the bar. Transverse reinforcement should be instrumented as well. Surface concrete strains should be monitored in the test region using strain gages and/or Zurich- type gages by means of a reference grid attached to the surface of the specimen. 2.5 Summary Based on the results of the literature review conducted in this chapter, the experimental plan for NCHRP Project 12-60 was refined and carried out with the approval from the proj- ect panel. The results of the entire experimental program for both strand and mild steel are presented and discussed in Chapter 3 of this report.

Next: Chapter 3 - Experimental Program and Results »
Transfer, Development, and Splice Length for Strand/Reinforcement in High-Strength Concrete Get This Book
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TRB's National Cooperative Highway Research Program (NCHRP) Report 603: Transfer, Development, and Splice Length for Strand/Reinforcement in High-Strength Concrete explores recommended revisions to the American Association of State Highway and Transportation Officials Load and Resistance Factor Design (LRFD) Bridge Design Specifications, which are designed to extend the applicability of the transfer, development, and splice length provisions for prestressed and non-prestressed concrete members to concrete strengths greater than 10 ksi.

Appendices A and B are published as part of NCHRP Report 603. Appendices C through I are available online via the links below:

* Appendix C: Rectangular Beam Summaries-Strand D

* Appendix D: Rectangular Beam Summaries-Strands A&B

* Appendix E: Rectangular Beam Summaries-Strand A (0.6 in.)

* Appendix F: I-Beam Summaries-0.5-in. Strand

* Appendix G: I-Beam Summaries-0.6-in. Strand

* Appendix H: AASHTO Mxxx-Standard Test Method for the Bond of Prestressing Strands

* Appendix I: NASP Test Protocols

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