Reactive Powder Concrete Anchorage for Post-Tensioning with Carbon Fiber-Reinforced Polymer...

(ProQuest Information and Learning: ... denotes formulae omitted.)

INTRODUCTION

Structures and structural components are often prestressed to improve structural efficiency and strength. Steel tendons have been used to induce the prestress. There are numerous problems with the corrosion

of steel tendons, however, leading to costly repair and refurbishment. Advanced composite materials (ACM) offer a distinct advantage in their durability and strength compared with steel. Recently, fiber-reinforced polymer (FRP) tendons have been introduced into reinforcing and prestressing of concrete structures to replace steel tendons. Fiber-reinforced polymer (FRP) prestressing tendons are stronger and lighter than their steel counterparts.1 FRPs will allow considerable development for sustainability and lower maintenance costs in construction, particularly in the fields of concrete and masonry, where corrosion problems could be eliminated.1-5 Hence, considerable effort has been expended to develop carbon fiber-reinforced polymer (CFRP) tendons into a corrosionfree prestressing system. One major obstacle to overcome has been the development of a simple nonmetallic anchor for these tendons for use on site.

A prototype nonmetallic anchorage for CFRP tendons was developed previously from a specially developed ultra high-performance concrete (UHPC).6-8 UHPC strengths were in the order of 200 MPa (29 ksi) at 7 days.9-16 Microcarbon fibers with a length of 3 mm (0.12 in.) were incorporated in the mixture to increase fracture toughness. The anchorage consisted of a barrel and four wedges all made out of the UHPC. The anchor barrel was wrapped in CFRP sheets to provide the tensile strength to resist barrel failure when the tendon is loaded and the wedges engage in the barrel. The problem with this nonmetallic anchor is its large size compared with a standard steel prestressing anchorage (120 mm [4.7 in.] diameter by 180 mm [7.1 in.] length). A smaller, reliable nonmetallic anchorage needed to be developed to promote the use of FRP and nonmetallic anchorage in concrete and masonry applications. Hence, the authors developed a reactive powder concrete (RPC) mixture (Table 1) with post-casting treatment procedures (applying pressure during setting and post-set thermal regimes to obtain a high strength, tough concrete material).10-13 This material was used to make new, small anchors for FRP tendons.

RESEARCH SIGNIFICANCE

The objective of the research described was to develop a new carbon fiber-reinforced reactive powder concrete (CFRRPC) anchorage for use with CFRP prestressing tendons. The new anchorage should be approximately the size of a standard steel anchor, durable, economic to produce, and easy to assemble and use on site. The anchorage should successfully grip FRP tendons mechanically to provide a metal-free, noncorroding prestressing system.

Initial development of a metallic anchorage and efforts to develop a UHPC nonmetallic anchorage showed the essential needs for a material of very high toughness for the wedges and a material with high tensile strength for the outer barrel. Whereas different materials could be used for the two components, such as ceramic (determined to be too expensive) or plastics (expectation of excessive creep), it was deemed appropriate to develop a new durable concrete for this purpose.

RPC is a recent development in concrete technology, with a wide scope of application opportunities, which could have significant impact on construction and structural engineering. It was decided to develop a tough RPC for the anchorage. The characteristics of the RPC developed permitted a reduction in anchorage dimensions and the number of CFRP wraps required for wrapping the barrel. The resulting system passed the PTI requirements for a prestressing/posttensioning system, thus offering, for the first time, the potential of corrosion-free prestressing of structures.

New CFRRPC anchorage system

The new anchor was designed for use with 8 mm (0.3 in.) diameter indented spiral CFRP tendons. The mechanical and physical properties of the tendons are given in Table 2. The anchorage consists of an outer casing (barrel) and four piece wedges, shown schematically in Fig. 1. The conical hole of the barrel is inclined at an angle ?^sub 1^ = 2 degrees while the wedges outer surfaces are inclined at an angle ?^sub 2^ = 2.1 degrees. The differential angle concept was based on the results from a metallic anchorage6-8 to give more even gripping pressure on the tendon than occurs if the angle of the wedges and barrel are the same.

Experimental testing of the tendons using the previously developed steel anchorage (80 mm [3.2 in.] long by 50.8 mm [2 in.] diameter) showed that the ultimate strength of the tendon varied from 105 to 124 kN (23.6 to 27.9 kip).13 These loads are slightly higher than the nominal failure strength of 104 kN (23.4 kip), quoted by the manufacturer, used in the design of the new CFRRPC anchorage system.

The anchor barrel is wrapped in CFRP sheets to provide the tensile strength in the hoop direction to resist cracking of the barrel radially when the wedges grip in the barrel. Smooth surfaces are needed for the barrel-wedge interface to decrease the friction force between the barrel and wedges. Low friction forces will reduce the tensile stresses in the barrel to ease the sliding of the wedges over the barrel's inner surface as the wedges seat. Initial seating of the wedges is needed to achieve grip on the tendon prior to applying load to the FRP tendon. Seating is performed with a hydraulic jack.

Design of CFRRPC anchorage system

Static model analysis-Preliminary anchorage dimensions were obtained using the static model created by Campbell et al.6 A schematic of the static model is shown in Fig. 2. From the free-body diagram, the following equilibrium conditions can be applied

(Tendon)F^sub TW^ = P (1)

(Wedges)F^sub TW^ = R^sub WB^ sin?^sub 2^ + F^sub WB^ cos?^sub 2^ (2)

where ?^sub 2^ is the wedge angle, P is the ultimate tendon strength (104 kN [23.4 kip]), R^sub WB^ is the friction force between the wedges and the tendon, and R^sub WB^ and F^sub WB^ are the normal and friction forces between the barrel and the wedges respectively. Friction produces

F^sub WB^ = ?^sub WB^R^sub WB^ (3)

where ?^sub WB^ is the coefficient of friction between the wedges and the barrel. Using Eq. (1), (2), and (3)

... (4)

Substituting the angle ?^sub 2^ = 2.1 degrees and assuming the coefficient of friction ?^sub WB^ between the concrete wedges and the concrete barrel of 0.1 (the wedge and barrel surfaces are smooth and grease was applied between them), Eq. (4) provides

... (5)

An increase in ?^sub WB^ results in a nonlinear increase in R^sub WB^. Thus, keeping a low coefficient of friction between wedge and barrel contact surfaces results in low normal forces and stresses exerted by the wedges on the barrel.

The barrel was designed to resist the circumferential tensile stresses that develop to resist R^sub WB^, which could lead to splitting of the barrel. Due to the shape of the wedges, stresses exist between the wedges and the barrel. Assuming a linear distribution of the radial stresses exerted on the barrel by the wedges varying from zero at the bottom of the wedges to a maximum at the top of the barrel, the maximum radial stress s^sub WB^ is

... (6)

where l^sub BC^ is the wedge-barrel contact length and assumed to be 75 to 80% of the barrel total length l^sub B^, and d^sub B^ is the top inner barrel diameter (assumed to equal the diameter of the steel anchorage (50 mm [2 in.]). One can estimate the tangential stress in the barrel from thin cylinder theory, taking l^sub BC^ = l^sub B^, as

... (7)

... (8)

where t^sub Bt^ is the wall thickness at the top of the barrel. Calculating the tangential stress created in the barrel would show the stress in the concrete and if there is a need for FRP sheet reinforcement for the barrel.

For example, when seating of the RPC anchorage was attempted, failure occurred at a tendon force of 40 kN (9 kip), with splitting occurring at the top of the barrel.

Using P^sub max^ = 40 kN (9 kip) and l^sub B^ = 100 mm (4 in.) and r^sub i^ = 11.5 mm (0.45 in.), which is the top inner radius of the barrel, s^sub T^ = 8 MPa (1.2 ksi). As the tensile strength of the concrete was expected to be approximately 25 MPa (3.6 ksi), the thin-walled model is clearly inaccurate.

Thick-wall hollow cylinder analysis-Another estimate of the highest tangential stress in the outer barrel can be made using thick cylinder analysis (Fig. 3).

Radial stress, thick cylinder

... (9)

Circumferential stress, thick cylinder

... (10)

where the subscripts i and o refer to inner and outer sections of the barrel, respectively.

The maximum internal radial pressure acting on the internal barrel surface is as in Eq. (6).

Assuming that p^sub o^ = 0, r^sub i^ = 13.5 mm (0.53 in.), and r = r^sub i^, then

... (11)

Solving for s^sub ?max^ [implies] = s^sub ?max^ = 184 MPa (26.7 ksi), which shows that CFRP sheet wrapping is required to avoid barrel splitting. Final detailed dimensions of the anchorage components were taken to be as in Fig. 4.

Barrel dimensions-The bearing stress on the anchorage is calculated in Eq. (13) assuming that the full base of the anchorage is bearing on the reaction surface. The outer diameter of the anchorage is governed by the bearing strength of the CFRRPC.

... (12)

... (13)

For a tendon load of 70 kN (15.7 kip) (70% of ultimate), the bearing stress is 56 MPa (8.1 ksi), well within the strength of the CFRRPC.

CFRP sheet wrapping of barrel

CFRP sheets were needed as circumferential tensile reinforcement of the barrel in order to resist the barrel splitting when the tendon is loaded.19-22 The engineering properties of the CFRP sheets used21 are given in Table 3. Assuming the full tensile splitting load is to be carried by the FRP wrap, the number of sheets needed was calculated from22

... (14)

where t^sub sheets^ is the total thickness of the wrapping sheet, f^sub sheets^ is the material factor for CFRP, and f^sub sheets^ is the ultimate strength of the CFRP sheets. Because the tendon allowable load is P = 104 kN (23.4 kip), and the length of the barrel was taken to be to be l^sub B^ = 100 mm (4 in.) (as the steel anchor), l^sub W^ = 75 mm (3 in.), where l^sub W^ is the wedges length, the material factor f^sub sheets^ = 0.9, and the ultimate strength of CFRP sheets, f^sub sheets6 = 3800 MPa (551 ksi), then one sheet was required to wrap the barrel.

When a PTI test was conducted, however, splitting of the barrel occurred at Stage II of the PTI test; therefore, a second CFRP sheet was added.

EXPERIMENTAL RESULTS AND DISCUSSION

Casting and curing of anchorage components

Casting-Barrel and wedge mold components were machined from AISI 4340 steel. A schematic of the barrel and wedge steel molds are shown in Fig. 5(a), and a picture of the actual molds in Fig. 5(b).

Side and bottom bleeding grooves were incorporated in both molds to facilitate escape of excess water and air when applying the presetting load.

Each mold consisted of a circular base (25 mm [1 in.] thick), four pieces for the housing, a core, and a cylindrical pressure cap. The core of the wedges for the mold consisted of four 1 mm (0.04 in.) thick steel plates welded to an 8 mm (0.31 in.) diameter steel rod. These plates separated the cast CFRRPC wedges. Eight hardened steel drive pins ensure precise alignment of the different mold components, prior to bolting during assembly.

Fresh CFRRPC was very viscous, requiring long periods of consolidation using vibration. The molds were placed on the vibrating table and the CFRRPC mixture was added in small portions for a casting period of 10 to 15 minutes.

Application of presetting pressure

The specimens of fresh concrete cast in steel molds were then subjected to a pre- and during-setting pressure of 26 MPa (3.8 ksi) under load control conditions in a close-loop electro-hydraulic test machine for 6 to 12 hours. More concrete was added as needed to compensate for the settlement that occurred due to the escape of water and air, until the cap no longer displaced under the load.

After setting, the molds were taken apart, and the specimens were removed for heat-treatment. The CFRRPC anchorage components had shiny and smooth surfaces similar to buffed marble.

Heat-treatment

The components were placed in an autoclave oven for 3 days for curing. The target curing temperature was 150 ?C (302 ?F). The curing temperature was both raised and lowered gradually (30 ?C/hour [86 ?F/hour]), to avoid any thermal shock that might cause cracking.

Dimensions and weight were measured before and after heat-treatment. Dimensional changes were of interest because they could alter the differential angle between the barrel and the wedges, which could have led to anchorage seating problems and/or tendon slipping. Whereas no dimensional changes were measured, some weight loss did occur. The weight loss was approximately 3 g/hour (0.12 oz/hour), directly proportional to the heat-treating temperature, and heat-treatment time (up to 6 hours).

Considerable care was taken during casting, application of pressure, and heat treatment to achieve consistency throughout specimen preparation.

The molds were cleaned thoroughly after every cast using a metal solvent cleaning liquid and scrubbed with a brass brush to remove debris, particles, or dust that might have interfered with the functionality of the bolts when reassembling the mold.

A CFRRPC barrel and wedges are shown in Fig. 6, together with a steel barrel and steel wedges for size comparison.

Wrapping of barrel

The mating ends of the wrapping sheet were not overlapped but butted and covered with second CFRP sheet (one could use a 25.4 mm [1 in.] wide CFRP sheet). Overlapping results in a void that may lead to barrel failure by ripping in the fiber wrap or splitting of the barrel material when the tendon is loaded. The CFRP sheet was cut to the barrel's outer area dimensions. Because the butt joint does not provide a complete circumferential wrap, a second sheet was placed over the first, with its butt joint diametrically opposite that of the first sheet. Thus, a continuous wrap around the barrel was provided. The direction of the carbon fibers in the CFRP sheets was perpendicular to the barrel axis (circumferential around the barrel). Barrels were wrapped after the CFRRPC anchorage elements had been cured (4 days). Barrel wrapping involved three tasks: application of the resin, wrapping, and pressurizing. (Note that due to the low permeability of CFRRPC, application of the primer material proved unsuccessful and unnecessary. Therefore, the use of the primer in wrapping the CFRRPC barrel was excluded.)

Pressure was applied immediately after wrapping, during saturant (resin) curing, to ensure excellent adhesion and compaction of the CFRP sheets against the barrel and successive CFRP sheets and to remove air bubbles entrapped during rolling. Application of pressure was also essential to avoid wrinkling of the CFRP sheets. Pressure was applied in two stages: 1) through a latex membrane vacuum chamber; and 2) with a desiccating bell.

For the vacuum chamber, a latex tube of 44.45 mm (1.75 in.) diameter and 177.8 mm (7 in.) long membrane was stretched over the upper and lower chamber edges, a vacuum was then applied to stretch the membrane open. The CFRP-wrapped barrel encapsulated in waxed paper was then placed in the center cavity of the latex membrane. The vacuum was then removed and the membrane contracted to create uniform pressure around the CFRP wrapped barrel. Unwrapped and CFRP-wrapped CFRRPC barrels are shown in Fig. 7.

A desiccating bell was subsequently used to ensure excellent bonding of the CFRP wraps to the barrel and complete the removal of any air bubbles or voids. Specimens enclosed in their latex membranes were placed in a glass desiccating bell and the vacuum-sealed cover replaced. A vacuum was then applied -762 mm Hg (-30 in. Hg) for approximately 1 day (the resin saturant cures completely in approximately 16 hours). The vacuum sucks the wrap onto the surface of the specimen as the membrane presses on the wrap as well. The wrapped barrel was then taken out of the desiccating bell and the latex and wax paper removed. Light grinding of both ends of the barrel was required in order to create flat and smooth surfaces. Grinding was performed using a sanding belt (60 grit sanding paper).

PTI test

Initially, the inner wedge surfaces were roughened and sharp edges removed using steel wool in order to increase friction between the tendon and the wedge surfaces to avoid slippage or premature rupture of the tendon under loading. In later tests, however, this sanding was omitted. Heavy-duty industrial grease (titanium grease or molly lubricant) was applied to the internal surface of the barrel to improve sliding of the wedges against the inner surface of the barrel. Prior to testing a tendon anchorage combination, the wedges were seated to provide initial grip of the tendon by the wedges.

The PTI requires a tendon-anchorage assembly to be tested under fluctuating cyclic loading (fatigue). The test is required to verify the durability and reliability of the anchorage system. CFRRPC anchorage systems were tested for compliance to the latest PTI requirements.23,24

The PTI-recommended test for tendon-anchorage assemblies involve three stages, applied according to the following sequence:

* Stage I: 500,000 cycles fluctuating between 60 to 66% of the nominal tendon strength (62 to 68 kN [14 to 15.3 kip]) for 27.8 hours, at five cycles/second;

* Stage II: 50 cycles fluctuating between 50 to 80% of the nominal tendon strength (52 to 83 kN [11.7 to 18.7 kip]) at one cycle/second; and

* Stage III: A static tension test up to 95% of the nominal tendon strength (99 kN [22.3 kip]).

The tensile strains induced in the CFRP sheets wrapping the barrel were monitored to understand the anchorage behavior as the tendon was loaded.

Data (strain in tendon and barrel CFRP wrap, load, and stroke) were collected through a computer at a sampling rate of 20 Hz for periods of 10 seconds at 30 minute time intervals.

The strain in the wrap around the barrel initially reduces with increasing load in the tendon (Fig. 8). This is because the pre-seating of the wedges causes outwards pressure on the barrel, resisted by circumferential tension in the wrap. As the tendon is loaded, not only does it lengthen, but it also gets thinner due to Poisson's effect. Thus, the pressures between tendon and wedges, and wedges and barrel are slightly relieved. Concomitantly, the circumferential tension in the wrap reduces. What is shown in Fig. 8 is the change in strain in the wrap from the preseated condition. Once the seating load (70 kN [15.7 kip]) is exceeded, the tendon can slip or, as occurs here, the tendons can pull the wedges more tightly into the barrel, increasing the pressure between wedges and barrel further. The result is increasing strain in the wrap.

PTI Stage I

Tendon slippage was experienced in the first test, so a seating force of 70 kN (15.7 kip) was applied for the remaining three PTI tests, with no subsequent slippage.

The relationship between tendon load and stroke at various numbers of cycles during the second PTI fatigue test for Specimen 2 is shown in Fig. 8.

Increasing tendon load resulted in a decrease in the tensile strain in the CFRP sheets, especially as the anchorage was being seated. The tensile strain in the barrel increased after the seating force of 70 kN (15.7 kip) was exceeded. Using a seating force of 70 kN (15.7 kip) resulted in less change in the measured tensile strain in the CFRP sheets in test numbers 2, 3, and 4, with no slippage of CFRP tendon. There were marginal fluctuations in the tensile strain in the CFRP sheets as the load reached its maximum level during the first few cycles. There is little subsequent strain change as the cyclic loading continued. This can also be related to further seating of the anchorage. The cyclic loading had no effect on the stiffness of the CFRP tendon tested.

PTI Stage II

In Stage II of the PTI fatigue test, again there was no effect on the stiffness of the tendon-anchorage assembly, monitored through the load/stroke relation throughout the whole test, nor on the CFRP tendon stiffness. When the anchorage was loaded to 83 kN (18.7 kip), the tensile strain in the CFRP sheets wrapping the barrel reached its maximum value, followed by some marginal fluctuations in the subsequent cycles. Further application of load resulted in an increase strain in the CFRP wrap as shown in Fig. 9.

In all fatigue tests (PTI Stages I and II), the tensile stresses in the CFRP sheets did not exceed 10% of their nominal strength.

PTI Stage III

Stage III is the final required PTI tension test involving a static tension test up to 95% of the nominal tendon strength (99 kN [22.3 kip]). In Test 2, the tension test was performed up to the failure load on the CFRP tendon (103 kN [23.2 kip]). The capacity of the MTS machine was exceeded by increasing the pressure in the hydraulic pump. Tests 2 and 3 were carried out to the maximum capacity of the MTS machine with no failure of the tendons at loads less than 100 kN (22.5 kip).

Results from the tests are listed in Table 4. An average anchorage efficiency ? of 96.4% was obtained. Anchorage efficiency ? was calculated using the following equation18

... (15)

where F^sub tu^ is the failure force of the tendon, A^sub sp^ is the cross sectional area of the tendon, and f^sub tu^ is the nominal failure force of the tendon.

The CFRP tendons showed an average elastic modulus of 165 GPa (24,000 ksi). Average ultimate strains of 1.28, 1.25, and 1.3% were observed in test numbers 1, 2, and 3, respectively.

The tensile stresses in the CFRP sheets around the barrel were monitored in order to study the behavior of the barrel during tendon loading. When the tendon load exceeded 90% of nominal tendon strength, the tensile stresses in the CFRP sheets did not exceed 1.5% of the nominal strength of the sheets (assuming an E = 227 GPa [33,000 ksi] and linear elasticity). Minimal variations in tensile stresses in the CFRP sheets were noticed in all the tests performed. This is especially noticeable in Stage III (tension test) up to a load of approximately 85 kN (19 kip), which can be attributed to the tensioning of the CFRP sheets during the pre-seating prior to the fatigue tests (Stage I and Stage II), as shown in Fig. 10. But once the pre-seating load was exceeded, the tensile stresses in the CFRP sheets increased to reach a maximum of approximately 50 MPa (7.3 ksi) at tendon failure. These stresses are 3% of the nominal strength of the CFRP sheets.

In Tests 2, 3, and 4, no cracks were noticed in the barrel even when the tension test was carried out to failure of the CFRP tendon. The wedges were therefore punched out of the barrel after testing using a hydraulic or manual punch in order to observe the inner parts of the anchorage. Some cracks were noticed on the wedges, but of more interest is the fact that the wedges were welded together due to the pressure during seating and testing. The barrels can be reused with new sets of wedges.

CONCLUSIONS

A new non-metallic CFRRPC anchorage system was designed, implemented, and tested against PTI requirements. The system complied with the required efficiency and fatigue resistance for prestressing and post-tensioning applications.

The new CFRRPC anchorage system will provide a completely metal-free environment, with similar dimensions to the previously developed steel anchorage (50 mm [2 in.] diameter by 100 mm [4 in.] long for CFRRPC anchorage and 80.8 mm [3.2 in.] diameter by 80 mm [3.1 in.] long for steel anchorage) and smaller dimensions, than the previously developed UHPC anchorage.

The weight of the new CFRRPC is approximately onethird that of the steel anchorage. The average weight of the new CFRRPC anchorage systems, including barrel, wedges, and CFRP wrapping is 475 g (16.8 oz), whereas the average weight of the steel anchorage is 1300 g (2.9 lb).

The use of RPC as a material provides a corrosion-free environment and reduces both the dimensions and the cost of the new non-metallic CFRRPC anchorage system.