105-s46

478 ACI Structural Journal/July-August 2008

ACI Structural Journal, V. 105, No. 4, July-August 2008.MS No. S-2007-058 received February 7, 2007, and reviewed under Institute publication

policies. Copyright © 2008, American Concrete Institute. All rights reserved, including themaking of copies unless permission is obtained from the copyright proprietors. Pertinentdiscussion including author’s closure, if any, will be published in the May-June 2009ACI Structural Journal if the discussion is received by January 1, 2009.

ACI STRUCTURAL JOURNAL TECHNICAL PAPER

The behavior of offset mechanical splices is investigated in thecontext of currently-prescribed acceptance test methods andcriteria. Two commercially available offset mechanical splicesystems were tested in direct tension with the splice both restrainedand unrestrained from rotating. In this series of tests, thecommonly held belief that the critical case is where the splice isnot restrained from rotating is shown to be incorrect. The splicesare additionally tested in air under prescribed fatigue loadingconditions. Finally, the splices are tested in place in 4.7 m (15 ft5 in.) long concrete beams under both monotonic and fatigueloading conditions and their behavior is assessed. Significant findingsinclude a classification of splice failure modes and characterization ofin-place splice behavior.

Keywords: acceptance test; reinforcement bar; splice.

INTRODUCTIONThe behavior of offset mechanical splices is investigated

in the context of currently-prescribed acceptance testmethods and criteria. Two commercially available offsetmechanical splice systems were tested in direct tension withthe splice both restrained and unrestrained from rotating. Inthis series of tests, the commonly held belief that the criticalcase is where the splice is not restrained from rotating isshown to be incorrect. The splices are additionally tested inair under prescribed fatigue loading conditions. Finally, thesplices are tested in place in 4.7 m (15 ft 5 in.) long concretebeams under both monotonic and fatigue loading conditionsand their behavior is assessed. Significant findings include aclassification of splice failure modes and characterization ofin-place splice behavior.

RESEARCH SIGNIFICANCEThe objective of this work was to assess the applicability

of performance criteria by which mechanical splices areassessed in the context of available offset mechanical splicesystems. It will be shown that conventional wisdom associatedwith testing offset splices is apparently incorrect. The splicesystem performance was also evaluated both in terms ofacceptance-type tests and in-place applications. Significantfindings include a classification of splice failure modes andcharacterization of in-place splice behavior. The presentwork is believed to be the only recent study of its kind.

REINFORCEMENT BAR SPLICESReinforcement bars are most often spliced in place using

lap splices. Lap splices place two bars adjacent to each otherover a sufficient length to affect full development of bothbars through stress transferred through the surroundingconcrete. The typical required length for a tension lap spliceis on the order of 50 to 70 times the diameter of the barsbeing spliced (ACI Committee 318 2005). Lap splices are

not permitted for bars larger than No. 11 (ACI Committee318 2005) and are often impractical, regardless of the barsize, in many applications. Alternatives to lap splices includewelded connections or mechanical connections.

Mechanical connections are divided into two categoriesbased on the expected mechanical loading applied to thesplice (ACI Committee 439 2007). Type 1 splices are usedwhen there is no expectation of inelastic deformation orelevated tensile stress due to seismic loading. Type 2 splicesare those that have been demonstrated through acceptedtesting procedures to be able to develop the specified tensilestrength of the spliced reinforcing bars for resistance toincreased tensile forces that may be expected from seismicloading. The use of Type 2 mechanical splices is referred toonly in the seismic provisions of ACI 318-05 (ACICommittee 318 2005), whereas Type 1 mechanical splicesare addressed in the body of the Code. Proposed revisions ofACI 439.3R (ACI Committee 439 2007) recommend the useof Type 2 mechanical splices over conventional lap spliceswhere inelastic yielding may be experienced. This recommen-dation is based on the observation that lap splices typically donot perform well under inelastic yielding conditions.

There are many situations that require the use of mechanicalsplices over the use of conventional lap splices. Mechanicalsplices are an attractive alternative for providing continuityand anchorage to hoop or continuous spiral reinforcementused to provide confinement in columns. Other applicationsinclude relieving congestion and reducing the reinforcementratio in splice regions and in splicing new reinforcing steel toexisting steel in patches, closure pours, and structural additions.Current codes do not permit No. 14 or No. 18 bars to bespliced using a lap splice, requiring mechanical splices forthese bar sizes. Other uses of mechanical splices are inportions of a structure affected by seismic loads as recom-mended by revisions to ACI 439.3R (ACI Committee 4392007). Finally, in the case of epoxy-coated or lower tensilestrength reinforcing bars, mechanical splices may representa practical alternative to the relatively long lap splicesrequired in these cases.

There are many types of mechanical splicing productsavailable. In this discussion, they have been categorized as1) in-line splices, in which the centerline of each spliced barcoincides; and 2) offset splices (also referred to as an offsetmechanical splice or a mechanical lap splice), where thecenterlines have an eccentricity.

Title no. 105-S46

Experimental Study of Offset MechanicalLap Splice Behaviorby Keith L. Coogler, Kent A. Harries, and Marcella Gallick

479ACI Structural Journal/July-August 2008

MECHANICAL LAP SPLICE PRODUCTSCurrently, there are two mechanical lap-splicing products

available:1. Product A, shown in Fig. 1(a), is a sleeve that allows

two bars to be placed side by side. Allowing a length of barof at least one bar diameter to protrude from each end, thehardened, pointed set screws are tightened through the top ofthe sleeve securing the bars in place. The connection is acombination of mechanical (screws penetrating into reinforcingbar) and friction (far side of bar bearing against sleeve).When tightened to an appropriate torque, the screw head willshear off, indicating uniform tightening of all screws. Thissplice is designed to carry tension and compressive forcesbut is currently only recommended for tension use. A uniqueaspect of this system is that it may be used to splice bars ofdifferent sizes provided the bars are only one standard sizeremoved (No. 3/No. 4, No. 4/No. 5, No.5/No. 6, and No. 6/No. 7); and

2. Product B, shown in Fig. 1(b), is an oval-shaped sleevewith a wedge-shaped pin inserted into it. The reinforcingbars to be spliced are positioned inside the sleeve and thewedge is inserted using a proprietary hydraulic pin driver.The wedge drives the bars against the outer walls of thesleeve affecting a friction connection to hold the bars inplace. Additionally, the hardened wedge deforms the bar asit is driven resulting in a further mechanical connection. Thissplice is currently only recommended for tension use.

The splicing sleeve, in each case, is sized according to thebar size to be spliced. The physical size of the body of bothmechanical lap splice types result in reduced concrete cover.In the case of Product A, the clear cover to the splice is 20 mm(0.79 in.) less than that to the spliced bar. The smallerProduct B results in a 6 to 9 mm (0.24 to 0.35 in.) reductionin clear cover to the spliced bar.

PERFORMANCE SPECIFICATIONSThe performance of mechanical splice systems is evaluated

with different testing procedures and requirements varyingby specifying agency. Typical requirements are listed inTable 1. Some jurisdictions have other related requirements;for example, Oregon requires a completed mechanical spliceto achieve a tensile capacity of 1.35fy rather than the typical1.25fy required by others. It is noted that California TestCT670 (Caltrans 2004) is a test method and does not specificallyrecommend acceptance criteria. The performance criteriaassociated with CT670 (Caltrans 2004) are those applied byCaltrans.

Test methods for mechanical splicesAssessing the performance of mechanical bar splices is

difficult and only recently has there been a uniform specificationgoverning these tests. ASTM A1034 (ASTM International

2005) is a new standard addressing testing of mechanicalsplices. ASTM A1034 (ASTM International 2005) providesonly general testing methodologies and does not providespecific parameters (such as the load at which to measureslip or the stresses appropriate for cyclic testing) and doesnot quantify any testing acceptance criteria. ASTM A1034(ASTM International 2005) includes an additional parameter—low temperature testing—where any of the standard tests areadditionally conducted at a reduced ambient temperature(not specified).

Before the ASTM A1034 (ASTM International 2005) spec-ification was released in late 2004, the CT670 (Caltrans 2004)Test Method was the only specification to specifically addressthe testing of mechanical splices. The CT670 (Caltrans 2004)test methods are outlined in Table 1 along with the acceptancecriteria typically associated with each test. Manufacturers ofoffset mechanical lap splice Types A and B report havingconducted direct tension testing of their product in the mannerdirected by CT670 (Caltrans 2004). Both products areapproved for use by Caltrans for Type 1 splices only.

Direct tension testing of offset spliced bar systems resultsin a moment being generated at the mechanical spliceresulting from the eccentricity of the bars. This moment willplace complex stresses on the coupler and result in thereinforcing steel kinking at or near the coupler face as theapplied tension loads try to align. This effect is shownschematically in Fig. 2.

REVIEW OF PREVIOUS STUDIESThe body of work addressing mechanical lap splices is

very limited with the only published work conducted by

Keith L. Coogler is a Structural Engineer at Westinghouse Electric Company,Monroeville, PA. He received his MSCE from the University of Pittsburgh, Pittsburgh,PA, in 2006.

Kent A. Harries, FACI, is the William Kepler Whiteford Faculty Fellow and an AssistantProfessor of structural engineering and mechanics at the University of Pittsburgh. Hereceived his PhD from McGill University, Montreal, QC, Canada, in 1995. He isa member of ACI Committees 215, Fatigue of Concrete; 335, Composite andHybrid Structures; 408, Bond and Development of Reinforcement; 439, SteelReinforcement; 440, Fiber Reinforced Polymer Reinforcement; and E803, FacultyNetwork Coordinating Committee.

Marcella Gallick is the Principal of Rhea Engineers and Consultants, Inc., GibsoniaPA. She received her MSCE from the University of Pittsburgh.

Fig. 1—Offset mechanical reinforcing bar splice systems.

Fig. 2—Effects of direct tension loading on mechanical splice.


Paulson and Hanson (1989 and 1991). Paulson and Hanson(1989) provided a summary and review of fatigue data ofwelded and mechanically spliced reinforcing bars. Thissurvey of existing research focused solely on fatigue data. Itwas noted that, at the time, there was no specification thatcontained provisions for the evaluation of fatigue of reinforcingbar splices. Comparisons were made with AASHTO LRFD(2004) design specifications for fatigue of straight unsplicedbars. Paulson and Hanson (1991) showed that mechanicallyspliced reinforcing bars may have a shorter fatigue life,although the fatigue life varies greatly based on the type ofsplice considered. Nonetheless, Paulson and Hanson(1991) concluded that for the splice systems tested, a fatiguefracture would occur in the bar near the spliced region, not inthe splice itself. Thus, the splicing hardware was sufficientlystrong although it affected the spliced bars in a manneraffecting fatigue resistance.

Paulson and Hanson (1991) report an extensive study of thefatigue behavior of welded and mechanical splices. This studytested mechanical lap splices using in-place beam tests andopen air axial tension tests. The authors report two open air

tension tests conducted on a No. 5 Type B splice. The ultimatestress values observed are reported as 438 and 614 MPa (63.5and 89 ksi). The reinforcing bar fracture of the first specimenoccurred inside the splice at the wedge, the second specimenfractured just outside the splice. The first specimen did notachieve an ultimate capacity of 1.25fy, failing Criteria I ofTable 1. The authors state that due to the offset of the splicedreinforcing bars, axial open air tension tests may not reflect thebehavior of the splice embedded in concrete.

Open air fatigue tests were only conducted on in-linespliced bars. Fatigue tests of Type B offset splices wereperformed on bars embedded in concrete beams. The beamswere 2133 mm (84 in.) long, 152 mm (6 in.) wide, and 203 mm(8 in.) deep with a nominal effective depth d of 152 mm(6 in.). Each beam had a single No. 5 bar as the primary flexuralreinforcement and each specimen was tested in third-pointflexure. There was heavy shear reinforcement located in theshear span but none in the constant moment region where thesplice was located. The beams also included crack formers toinduce flexural cracking at each end of the coupler. Five ofthe nine reported fatigue-induced reinforcing bar rupturesinitiated at the junction of the splice wedge and the bar; theremaining four occurred immediately outside the splice.Stress versus number of cycles to failure (S-N) behavior ofthe specimens having spliced bars was notably degraded ascompared with the bare bars tested in direct tension fatigue(Paulson and Hanson 1991).

Paulson and Hanson (1991) attempted to establish limits fordesign stress ranges appropriate for different classes ofsplices. The AASHTO (2004) specified limit for the fatiguestress range on straight reinforcing bars for service loads was138 MPa (20 ksi). The authors classified mechanical connec-tions into three categories assigning maximum allowablestress ranges of 28, 83, and 124 MPa (4, 12, and 18 ksi) as indi-cated in Table 2. All coupling methods were assumed toreduce the fatigue limit of the reinforcing bars to some degree.Based on the limited testing, Type B offset mechanicalcouplers were assigned the same category as swaged andthreaded in-line couplers, having a maximum allowablefatigue stress range of 83 MPa (12 ksi). Couplers in this groupare characterized as causing a reduction in area of the splicedbar as results from machining threads, swaging the bar orinstalling the wedge. The results obtained by Paulson andHanson (1991), although limited, point to a significantdifference in behavior between offset mechanical couplerstested in air and those tested in place in concrete beams.

EXPERIMENTAL PROGRAMThe two commercially-available offset mechanical splice

systems, previously designated Types A and B, were evaluatedin four series of tests. In each series (except the beam tests),at least five samples of each bar size (No. 4, No. 5, andNo. 6) were tested. The performance of each specimenwas evaluated in accordance with Criteria I-IV given inTable 1. The experimentally determined material propertiesfor the reinforcing steel used in all tests in this study arelisted in Table 3. The tests conducted on each splice systemare shown schematically in Fig. 3 and described as follows.

Direct tensionDirect tension (DT) tested the reinforcement bar splice in

open-air direct tension and allowed the splice to freely rotate(Fig. 2). Each specimen was instrumented to record the slipcomponent over the spliced region. The specimens were

Table 1—Mechanical reinforcing bar splice performance criteria

Standard

Criteria Descriptor Performance1 2 3 4 5

• • • • •* Criteria I: tensile strength of mechanical coupler > 1.25fy

• • •Criteria II: allowable slip resulting from applied stress of 0.50fy then

relaxed to 0.05fy< 0.25 mm

(0.01 in.)

• Criteria III: tensile strength of mechanical coupler > 0.90fu

•Criteria IV: allowable slip resulting

from +172 MPa to –172 MPa(±25 ksi) for 10,000 cycles

< 1.25 mm (0.05 in.)

•Criteria V: allowable slip resulting from cycling between 0.90fy and

0.05fy for 100 cycles< 1.25 mm

(0.05 in.)

*Tensile strength of mechanical coupler > 1.20fy .Notes: Standards cited: ACI 318-05 (ACI Committee 318 2005), AASHTO LRFD(2004), AASHTO ASD (1996), CT 670 (Caltrans 2004), and CSA S6-00 (CSAInternational 2000).

Table 2—Splice categories according to Paulson and Hanson (1991)Maximum stress level

28 MPa(4 ksi)

83 MPa(12 ksi)

124 MPa(18 ksi)

138 MPa(20 ksi)

Splice typeAll

welded splices

Cold swaged steel coupling sleeves; taper- and straight-

threaded steel couplers; steel cou-pling sleeve with wedge (Type B)

Grout- and steel-filled coupling sleeves

Straight bar (no splice)

Table 3—Reinforcing bar material propertiesNominal value* No. 4 No. 5 No. 6

Yield strength fy414 MPa(60 ksi)

448 MPa(65 ksi)

414 MPa(60 ksi)

414 MPa(60 ksi)

Yield strain εy 0.002 0.003 0.003 0.0035

Tensile strength fu621 MPa(90 ksi)

717 MPa(104 ksi)

690 MPa(100 ksi)

662 MPa(96 ksi)

Elongation at rupture 11%† 19%‡ 19%‡ 23%‡

*AASHTO M31 (1996).†Elongation calculated over 203 mm (8 in.).‡Elongation calculated over 76 mm (3 in.).

ACI Structural Journal/July-August 2008 481

loaded monotonically at a rate of approximately 900 N/second(200 lb/second) until rupture of the bar occurred or therecorded slip exceeded 25.4 mm (1 in.).

Restrained tensionRestrained tension (RT) tested the reinforcement bar splice

in a manner similar to a pullout test that inhibited the splicefrom rotating. In these tests, only one spliced bar was subjectto tension with the reaction being provided by the coupleritself. The uniform reaction between the coupler and loadingplate effectively restrained the tendency of the coupler torotate. To ensure that the coupler was correctly engaged, adummy bar was provided to mimic the bar engaged in theuntested side of the splice. Each of the specimens was loadedmonotonically to failure similar to the DT tests and the slip ofthe loaded bar through the splice was measured.

Fatigue testsFatigue (F) tests were modeled after the CT670 (Caltrans

2004) test method, which requires cycling the specimenthrough a 345 MPa (50 ksi) stress range. The minimumrequired fatigue life of 10,000 cycles could not be attainedunder the specified test stress range for these specimen types.The stress range was reduced to 138 MPa (20 ksi) and theload was cycled from 69 MPa (10 ksi) compression to 69 MPa(10 ksi) tension for 10,000 cycles. The test setup was thesame as the DT tests and thus the splice was free to rotate, asshown schematically in Fig. 2. Slip through the splice wasrecorded following cycling.

Flexural beam testsFlexural beam (B) tests were conducted with the reinforcement

bar splice embedded in concrete. Eight reinforced concretebeams were cast for this testing phase. Each specimen was254 mm (10 in.) deep, 305 mm (12 in.) wide, and 4743 mm(187 in.) long. Each beam had a single No. 4 reinforcementbar as the primary flexural reinforcement. All beams(monotonic and fatigue) were loaded in four-point flexureover a 4540 mm (179 in.) simple span. The reinforcing barsplice was located in the center of the constant momentregion. One beam of each pair was tested monotonically tofailure. The second beam of each pair was subjected to10,000 cycles of repeated loading intended to result in anapplied stress range in the No. 4 flexural reinforcement barof 138 MPa (20 ksi), similar to the FT tests. Companion testsof unspliced straight bars and conventional lapped bar spliceswere included in this series as described in the following.

FAILURE MODESFor consistency in reporting, four failure modes were identi-

fied and denoted A through D, as shown in Fig. 4. Thesefailure modes were only recorded in the DT and RT tests.Failure Mode A was a rupture of the reinforcing bar at a signif-icant distance from, and apparently unaffected by, the splice,similar to a straight bar test. Failure Mode B was a rupture ofthe reinforcing bar at the wedge or first bolt; this failureresulted from the stress raiser induced by the wedge or bolt.Failure Mode C was a rupture of the reinforcing barlocated just outside of splice caused by the kinking of the barat this location. Failure Mode D did not result in a rupturedbar but rather the bar slipping through the splice a distancegreater than one rib spacing. This failure mode resulted in agouge on the spliced bar resulting from the bolts or wedge asshown in Fig. 4. In the experimental program, Failure Mode D

was allowed to progress until the resulting slip exceeded25 mm (1 in.).

Slip measurementsTo assess bar slip Criteria II and IV (Table 1) in this test

program, the slip was measured using displacement transducersinstalled on the unloaded projection of the spliced bar (as shownin Fig. 3). In this manner, the recorded displacementincluded only the slip component over the spliced region anddid not include elastic or inelastic deformation in the bar.The total slip for an individual specimen was the sum of theslip of each bar. In this work, the slip reported in subsequentsections and in Table 4 is calculated as twice the maximumsingle bar slip observed in each specimen. This case is anupper bound, where both bars experience the samemaximum slip. Reporting in this manner also permits directcomparison between the DT and F tests and the RT tests.

RESULTS OF TENSION TESTSA summary of the results of DT, RT, and F tests is

provided in Table 4. In this table, average values associated

Fig. 3—Schematic illustrations and photographs of test setupsused. (Note: 1 mm = 0.039 in.)

Fig. 4—Failure modes of offset mechanical splices tested indirect or restrained tension.


with the splice assessment Criteria I through IV (Table 1) arepresented. Additionally, the number of tests passing thecriteria and the failure modes are indicated.

DT behavior—Representative stress-versus-bar-strainresults for the DT tests with No. 4 bars are shown in Fig. 5(a).The strains were recorded from electrical resistance straingauges mounted approximately 50 mm (2 in.) from the splicelocation. The stress-strain relationship had an initial linearelastic portion, which is predictable and repeatable at lowstress ranges. At an applied stress of 138 to 207 MPa (20 to30 ksi) for each specimen, there tends to be softening of thesystem indicated by a decrease in the slope of thestress-strain curve. This apparent change in stiffness iscaused by the rotation of the splice—the system sustainsincreasing loads but the reinforcing bar is no longer only inaxial tension. Similarly, stress-versus-slip relationships(Fig. 5(b)) are also predictable and repeatable. Thestress-slip relationship is generally linear until yield of thereinforcing bar, near 414 MPa (60 ksi). Following yield, theslip begins to increase greatly, often to more than 6 mm(0.24 in.) before eventual bar rupture.

As indicated in Table 4, although all but one specimenpassed both Criteria I and II, the capacity of the splices (asmeasured by bar stress developed) is diminished as the barsize increases. The Type A splices exhibited superiorcapacity, satisfactory slip values, and significantly lowervariance as compared with Type B splices. Failure Mode Cwas dominant and resulted from the kinking of the bar andthe eventual bearing of the kinked bar against the mechanicalsplice. This was more pronounced in smaller bar sizesbecause the flexural stiffness of the bar (and thus resistanceto kinking) is reduced.

RT behavior—There is an apparent consensus amongmanufacturers and users of mechanical splices that due to thekinking of the bar, the results of DT tests are conservative;that is, if a splice passes a DT test, it will perform better inplace where rotation may be restrained (Coogler et al. 2006).The RT series of tests was formulated to test this hypothesis.As indicated in the results presented in Table 4, the capacitymeasured using an RT test is lower than a comparable DT

Table 4—Average performance of offset mechanical splice systems tested in air

Test Performance criteria

Splice Type A Splice Type B

No. 4 No. 5 No. 6 No. 4 No. 5 No. 6

DT

Ultimate stress fu ,exp, MPaCriteria I and II:

fu > 1.25fy = 518 MPafu > 0.90fu = 559 MPa

Average 704.7 681.2 641.2 669.5 622.6 602.6

Standard deviation 0.5% 2% 1% 6% 5% 5%

Samples passing 5/5 5/5 5/5 5/5 5/5 5/5

Slip at 200 MPa, mmCriteria II:

Slip < 0.25 mm

Average 0.060 0.096 0.112 0.046 0.026 0.050

Standard deviation 58% 58% 78% 133% 60% 40%

Samples passing 5/5 5/5 5/5 5/5 5/5 4/5

Observed failure modes(number of specimens) B/C/D 0/5/0 2/2/1 1/4/0 0/5/0 2/3/0 2/3/0

RT

Ultimate stress fu ,exp, MPaCriteria I and III:

fu > 1.25fy = 518 MPafu > 0.90fu = 559 MPa

Average 670.2 650.2 628.1 537.1 625.4 510.9


Samples passing 5/5 5/5 5/50.9fu: 1/51.25fy: 3/5 6/6

0.9fu: 0/51.25fy: 2/5

Slip at 200 MPa, mmCriteria II:

Slip < 0.25 mm

Average 0.026 0.040 0.030 0.010 0.020 0.036


Samples passing 5/5 5/5 5/5 5/5 6/6 5/5

Observed failure modes A/D 0/5 0/5 0/5 0/5 1/4 0/5

FSlip following 10,000 cycles, mm

Criteria IV:Slip < 1.25 mm

Average 0.406 0.812 1.220 0.508 1.016 2.286


Samples passing 5/5 4/5 4/5 5/5 3/5 1/5

Notes: Bold entries indicate average value did not pass criteria. 1 MPa = 145 psi; 1 mm = 0.0394 in.

Fig. 5—Representative results for direct tension tests ofNo. 4 bars. (Note: 1 MPa = 145 psi; 1 mm = 0.039 in.)


test. The increased DT test capacity is attributed to thekinked bar bearing against the splice and engaging additionalfriction forces not present in the RT test. Slip valuesreported in the RT tests are reduced and the variance ofboth capacity and slip values are similar to that found inthe DT tests. The failure mode for the RT tests is almost exclu-sively Mode D, which is a behavior dominated by thenature of the mechanical connection.

Representative stress-versus-slip results for the RT testswith No. 4 bars are shown in Fig. 6. The stress-slip relationshipis predictable and repeatable. The stress-slip relationship isgenerally linear until yield of the reinforcing bar (414 MPa[60 ksi]). Following yield, the slip begins to increase greatly,often to more than 38 mm (1.5 in.). In each of the tests, thespecimen reaches an ultimate stress, the load begins todecrease as the reinforcing bar begins to be pulled throughthe splice, indicating a Failure Mode D. After a givendisplacement, the load increases and then decreases again.This effect was more pronounced for the Type B specimensbut was evident in nearly all of the RT tests. It was determinedthat the reinforcing bar ribs contributed to this apparentincrease in load-carrying capacity: as the bar slipped, the boltor wedge engaged subsequent reinforcing bar ribs, increasingthe pullout force. This conclusion is confirmed by the Type Bresults where the spacing, in terms of slip measurements, ofthe load increases corresponded to the bar rib spacing. Suchbehavior was less evident in the Type A splice as the multiplebolts gouged into the bar as the bar slipped (Fig. 4).

Similar to the DT tests, an increase in bar size resulted in adecrease in performance in the performance Criteria I and III(Table 1) for the Type A splice specimens as shown in Table4. The Type B specimens did not follow this trend andperformed marginally overall with respect to Criteria I and III.

The RT test setup used considered only pullout from one sideof the splice. While not reflecting in-place conditions, this setupovercame the need for large restraining forces and resulted inaccurate pullout capacities not affected by the kinking of the baror the binding of the bar along the edge of the splice thought toaffect the DT tests. Thus, for the purposes of product evaluation,it is felt that this simple test is appropriate.

F test protocol and behavior—Performance Criteria IV(Table 1) requires cycling through a stress ranging from172 MPa (25 ksi) in compression to 172 MPa (25 ksi) intension, a 345 MPa (50 ksi) stress range S for 10,000 cycles (N =10,000). Due to the stress raisers induced by the mechanicalsplices, the bolts, wedges, and the bar kinking at the couplerface, this stress range was unachievable if 10,000 cycles ofloading were required. Initial tests conducted at this stressrange resulted in fatigue-induced reinforcing bar ruptureoccurring at between 70 and 320 cycles for No. 4 and No. 5specimens of both splice Type A and B (Coogler et al. 2006).

Commonly accepted S-N relationships for straight,unspliced reinforcing steel (Helgason and Hanson 1974)predict a fatigue life of approximately N = 100,000 cyclescorresponding to a stress range of 345 MPa (50 ksi). It isexpected that mechanically-spliced bars having inherent stressraisers should have significantly reduced fatigue lives, espe-cially at higher stress ranges. A number of variations in thefatigue setup were attempted, including varying bar lengthsand stress ranges. Ultimately, the bar lengths had to be muchshorter than DT and RT tests to avoid buckling failures and theadditional stresses induced by buckling. Additionally, thestress range had to be reduced. The stress range selected forfatigue testing was 138 MPa (20 ksi), ranging from 69 MPa

(10 ksi) in compression to 69 MPa (10 ksi) in tension. Thisstress range: a) permitted 10,000 cycles to be achieved whilemaintaining a through-zero fatigue protocol felt to be criticalin assessing the behavior of the splices considered; b) exceedsthe AASHTO-permitted stress range for mechanical couplers;c) is equal to the AASHTO-permitted stress range for unsplicedreinforcing bars (and thus provides a point of comparison);and d) results in a similar stress range as that considered for theflexural beam (B) tests reported in the following.

Similar to the DT and RT tests, an increase in bar size resultedin a decrease in performance. For the Type A specimens, onlytwo specimens exhibiting a larger-than-allowable slip. TheType B specimens performed generally less well, exhibitingmore scatter with great variability in the results; six Type Bspecimens experienced slip values greater than allowable.

Flexural beam (B) testsTo assess the in-place performance of offset mechanical

splices, the splices of No. 4 bars were embedded in concretebeams and the beams were tested in two ways: 1) undermonotonically increasing load to failure; and 2) subject tofatigue-conditioning followed by a monotonically increasingload to failure. The behavior of the splice was assessed andcompared with the behavior of a straight unspliced No. 4 barand a standard 305 mm (12 in.) long lap splice.

Eight reinforced concrete beams were cast for this testprogram. Each specimen was 254 mm (10 in.) deep, 305 mm(12 in.) wide, and was simply supported over a 4540 mm(179 in.) span as shown in Fig. 3(d). Each beam had a singleNo. 4 reinforcing bar as the primary flexural reinforcementand two No. 3 bars in the compression zone. The beams werecast in pairs, two specimens each having: a) straight,unspliced bar (designated: C); b) standard 305 mm (12 in.)long conventional lapped bar splices (L) (compliant withACI 318-05 [ACI Committee 318 2005] and AASHTO[2004] ; c) Type A mechanical splices (A); and d) Type Bmechanical splices (B). One beam of each pair was testedmonotonically to failure (designated as indicated previously:C, L, A, and B). The second beam of each pair was subjectto 10,000 cycles of repeated loading intended to result in anapplied stress range in the No. 4 bar of 138 MPa (20 ksi) asmeasured in the first cycle of load, N = 1. The latter specimensare referred to as fatigue conditioned and are designated with atrailing F (that is, CF, LF, AF, and BF). Following fatigueconditioning, the specimens were loaded monotonically tofailure. The same No. 4 reinforcing steel used for open air

Fig. 6—Restrained tension stress-slip results for No. 4 bars.(Note: 1 MPa = 145 psi; 1 mm = 0.039 in.)


tests was used in this test series. The 28-day concretecompressive strength was 41.6 MPa (6030 psi) and all beamswere tested at ages from 86 to 121 days.

All beams were loaded in four-point flexure with a 910 mm(36 in.) constant moment region located in the center of a4540 mm (179 in.) simple span. The reinforcing bar splicewas located in the center of the constant moment region. Thesetup is shown in Fig. 3(d). Fatigue conditioning was appliedto the F beams in load control with the total applied load(actuator load) ranging from 3.6 to 9.8 kN (800 to 2200 lb)in a sinusoidal waveform having a frequency of 1.0 Hz. Theload range was selected based on the measured reinforcingbar strains from the monotonic test series. The 3.6 kN (800 lb)lower limit was selected to represent an applied dead load; the9.8 kN (2200 lb) upper limit was selected to develop thedesired strain and, therefore, stress levels in the No. 4 rein-forcing bar and across the splice. The target stress level was138 MPa (20 ksi), corresponding to the F tests describedpreviously. All data were reported in terms of total actuatorapplied load P. Thus, the resulting moment in the constantmoment region was 0.91P (kN-m) (2.98P [lb-ft]). Similarly, themaximum shear in the beam was 0.5P (kN or lb).

The monotonic tests and final cycle to failure for thefatigue tests were conducted in displacement control at a rateof 7.3 mm/minute (0.29 in./minute). Due to the stroke limitationsof the actuator and the ductility of the under-reinforced beams,additional spacers were required to test the beam specimensto failure. In each monotonic test, the beams were loaded toa deflection of 75 mm (3 in.), unloaded, and the resultingpermanent deflection was made up with spacer platesbetween the actuator and spreader beam. The test wascontinued to a deflection of approximately 127 mm (5 in.)where the beam came into contact with the test frame.Although not tested to their ultimate failure, the final deflectionswere all on the order of L/32 and thus may be reasonablyassumed to have exceeded the beams’ ultimate limit statesdemands. This load history described results in the loop at adisplacement of 75 mm (3 in.) evident in the presented load-deflection plots. During monotonic testing, the displacementswere held constant at specified load intervals to documentcracking and investigate the specimens’ behavior. The testswere paused less than 10 minutes in each case and the entiretesting time (to failure) was kept under 2 hours. It was determinedthat these pauses did not affect the behavior of the specimen.

Each beam was instrumented with electrical resistancestrain gauges on the No. 4 flexural reinforcement. Gaugeswere located 305 mm (12 in.) on each side of midspan andthus fell in the constant moment region. Vertical displacementswere recorded using draw wire transducers (DWT) located

under each load point (Fig. 3(d)). The hydraulic actuator wasequipped with an internal load cell.

Flexural beam test resultsThe load-versus-displacement results of each beam are

shown in Fig. 7. The behavior of the load-deflection plotsshow the ductility of each under-reinforced beam and indicateevidence of slip in the splicing methods or rotation of thesplice within the concrete as described further in thefollowing. The jagged behavior of these plots reflect therelaxation that occurred when the loads were held to assesscracking behavior. The displacement limitations of thetesting frame prevented testing of the specimens to theirultimate load-carrying capacity; thus, ultimate load cannotbe a basis of comparison between the different specimens.Therefore, applied load resulting in a specified deflection isused as a means of comparison. To compare the stiffness foreach specimen, the load and strain were recorded at specifieddisplacements and are presented in Table 5.

In Column 1 of Table 5, it is evident there was littledegradation of load-carrying behavior caused by thefatigue conditioning of each specimen type. Additionally,the control series (C and CF) exhibited the stiffest behavior,whereas the lap splice (L and LF) was the least stiff of allspecimens. There was little difference in stiffness betweenthe A (and AF) and B (and BF) specimens, although bothwere marginally less stiff than the control beams having acontinuous reinforcing bar. The behavior described indicatesa marginal reduction in capacity associated with each splice,which may be attributed to nominal slip or relative movement ofthe splice. Column 2 of Table 5 presents the strain values ata displacement of 50.8 mm (2 in.). From these values, it isevident there was some accumulated damage due to fatigueconditioning in all cases because the monotonic strain is lessthan the fatigue strain. The very large strains for the LF andAF specimens are likely caused by the presence of a flexuralcrack very near the gauge location.

Column 3 provides the applied load at a displacement of127 mm (5 in.) (near the maximum test setup deflection forall specimens). The fatigue conditioned control (CF) and thelap splice specimens (L and LF) performed in a similarmanner as the monotonic control Specimen C. For the Type Aspecimens, the fatigue conditioned specimen (AF) had ahigher load than the monotonic loaded specimen (A). Thisis explained by the fact that the monotonic Specimen Aclearly exhibited slip of the splice and began to shed loadas a result (described in detail in the following). The B andBF specimens performed similar to the A-series with thefatigue conditioned specimen achieving higher loads thanthe monotonic specimen. Again, marginal slip of the spliceduring the monotonic tests is believed to account for this asdiscussed in the following. The following discusses thebehavior of each test specimen.

Specimens C and CFSpecimen C had the highest recorded applied load of the

beam tests, which was a result of this specimen having ahigher post-yield stiffness than the other specimens. Thepeak load was near 20 kN (4500 lb), although this specimenwas loaded to a deflection closer to 152 mm (6 in.). Duringtesting of the other specimens, the tests were stopped atdeflections of approximately 140 mm (5.5 in.); the load inSpecimen C at this deflection value was 19 kN (4270 lb).Specimen CF showed little degradation from the fatigue

Table 5—Flexural beam test resultsat given displacements

Column 1 2 3

SpecimenLoad at 50.8 mm

displacement, kN (lb)Strain at 50.8 mm displacement, με

Load at 127 mm displacement, kN (lb)

C 15.1 (3390) 2217 19.0 (4270)

CF 15.2 (3420) 2262 18.1 (4070)

L 12.8 (2880) 2094 16.3 (3660)

LF 13.3 (2990) 9133 16.5 (3710)

A 15.0 (3370) 1282 11.3 (2540)

AF 13.9 (3120) 9015 17.5 (3930)

B 13.8 (3100) 1349 14.2 (3190)

BF 14.3 (3210) 2149 17.3 (3890)


conditioning, performing similar to Specimen C throughoutmost of the test history. Some degradation in the behavior ofSpecimen CF was evident at the end of the test at displacementsexceeding 127 mm (5 in.), although this behavior cannotnecessarily be attributed to the effects of fatigue conditioning.

Specimens L and LFSpecimens L and LF performed similarly with little difference

in stiffness due to the fatigue conditioning. This series had alower stiffness than the other specimens and a peak load ofonly 16.7 kN (3750 lb).

The apparent degradation of behavior of the L specimensas compared with the C specimens may be attributed to thesofter expected response of the lap splice as compared withthe continuous bar. In a conventional lap splice, relative slipof the bars, in addition to steel strain, contribute to themeasured elongation across the splice. The slip initiatesimmediately and increases until the bond stress is exhaustedat which point the lap splice can carry no additional load andeventually fails, shedding its load carrying capacity. The cyclicloading response of lap splices was observed to be significantlyinferior to the monotonic loading response. The bond stressesdeveloped in lap splices subject to cyclic loading histories wereobserved to deteriorate more rapidly than bond stresses undermonotonic loading (Viwathanatepa et al. 1979). Additionally,there is a general consensus (Viwathanatepa et al. 1979; Lukoseet al. 1982; and MacKay et al. 1988) that for cyclic loadingconditions, the effects of confinement reinforcement areinsignificant, although recent work by Harajli (2007)contradicts this. For the tests conducted in this study, notransverse confinement was provided and thus the deteriorationdue to cycling (or rather the beneficial effects of confinementunder monotonic conditions) was not evident.

Specimens A and AFSpecimen A performed well initially, reaching a peak load

of 15.6 kN (3510 lb). Upon reloading following holding atthis peak (to record cracking), however, the specimen neverregained its previous capacity, achieving a capacity of only14.2 kN (3190 lb) before the load began to decrease as thedeflection continued to increase, indicating a failure of thespecimen. Following testing, the splice was recovered andinspected. The splice exhibited clear signs of slip: one barslipped approximately 13 mm (0.5 in.) through the splice.

Specimen AF showed little sign of degradation resultingfrom fatigue conditioning and achieved a higher load thanthe monotonic test, reaching an ultimate load of 17.8 kN(4000 lb) at a deflection of 133 mm (5.24 in.).

In Specimens A and AF, the concrete was unable to properlyconfine the splice and there was cracking evident on the soffit ofthe specimen caused by the rotation of the splice or slip ofthe bars through the splice. This cracking demonstrated aparticular problem with offset splices: the cracking of thecover concrete may have caused particular problems instructural elements exposed to the environment, and especiallydeicing salts. This cracking is shown on Specimen A inFig. 8(a) at a load of 15.6 kN (3510 lb).

Fig. 8—Longitudinal cracking of beam soffits due to couplerrotation.

Fig. 7—Beam test load-deflection results. (Note: 1 MPa = 145 psi; 1 mm = 0.039 in.)


Specimens B and BFSpecimen B performed in a similar manner as Specimen A.

During the initial loading, the specimen had a reasonablestiffness. At the peak load of the initial loading step, thesplice began to rotate and the cover concrete began to spall.Figure 8(b) shows the effect of the splice rotation on thecover concrete. The cracking was first documented at 13.8 kN(3100 lb), and the cracking in Fig. 8(b) is shown at a load of15.1 kN (3390 lb) at the peak of the initial loading cycle.

During reloading, the specimen’s performance deterioratedand was unable to achieve higher loads as the deflectionscontinued to increase. Slip failure similar to that observed inSpecimen A was suspected. Upon post-test inspection,however, there was no noticeable slip, although a great dealof rotation occurred, which would also result in an increase indeflection without an increase in load. Similar to Specimens Aand AF, the concrete cover was unable to restrain thesplice from rotating.

Splice-induced damage to concreteSplice-induced damage was only significant following

yield of the beams. At service load levels, through all fatigueconditioning, and prior to yield, no splice-induced damage tothe concrete was apparent. As deformations increased into thepost-yield region, however, significant damage to the concretecover was evident. Figure 9 shows images of the beam soffitsfollowing testing. Figure 9(a) and (b) shows the expectedflexure-induced transverse cracking evident for Specimens Cand L. No other damage is apparent, including longitudinalcracking in Specimen L, which may indicate lap splice slip.The rotation and resulting loss of cover associated with eachmechanical splice is clearly shown in Fig. 9(c) though (f).

SUMMARY AND CONCLUSIONSTwo commercially available offset mechanical splice

systems, designated in this paper as splice Types A and B,were evaluated in four series of tests: DT, RT, F, and B tests.The performance of each specimen was evaluated in accordancewith a number of performance criteria indicated in Table 1.The following conclusions are drawn from this work:

1. An increase in reinforcement bar diameter from No. 4 toNo. 6 resulted in a decrease in performance for each ofthe criteria considered, although most specimens stillpassed the criteria;

2. Generally, Type A splices outperformed Type B splices;3. Contrary to manufacturers’ assumptions, the DT test was

not necessarily conservative; the splice capacity determined bythe DT test was greater than that determined by the RT testpresumably due to friction between the kinked bar and coupler;

4. Failure Mode C: The rupture of the bar at the stressraiser associated with contact between the kinked bar andcoupler was the most commonly observed failure mode inDT tests;

5. Failure Mode D: A pullout failure was the mostcommon failure mode observed for the RT tests. This modeof failure results in a decrease in apparent ultimate stress forthe system because of the inability to develop the fullstrength of the cross section;

6. A 345 MPa (50 ksi) stress range for F testing results infatigue-induced reinforcing bar rupture at a very low numberof cycles. A more reasonable stress range of 138 MPa (20 ksi) issuggested for assessing the performance of this type of splice;

7. There was no noticeable degradation of the in-placesplice behavior resulting from fatigue conditioning at a stressrange of 138 MPa (50 ksi) applied for 10,000 cycles; and

8. For all in-place testing, concrete was unable to properlyconfine the offset splice near ultimate load levels.

QUALITATIVE OBSERVATIONSAll mechanical splices were installed at the University of

Pittsburgh’s structural research laboratory following themanufacturer’s guidelines and specifications. As noted, theType B product requires the use of a proprietary hydraulicwedge driver and the Type A product can be installed usinga hand-held ratchet or torque wrench. The Type A productpresents more options if there are clearance issues wheninstalling; for example, a simple ratchet could be used toinstall the splice. The wedge driver requires time-consumingadjustments to the driver tool to splice different size bars,whereas the Type A product simply adds to the number ofscrews that need to be tightened—the screw head sizeremains constant.

A concern with both of the mechanical splices consideredherein are the dimensions of the product. The Type B spliceis much smaller and encroaches less on the amount of coverpresent when the splice is embedded in concrete. Nonetheless,little difference in concrete behavior was evident.

RECOMMENDATIONSThere is a limited body of knowledge on the testing and

use of offset mechanical splices. There needs to be furtherwork conducted in this area before the use of these splicescan be widely accepted. Some recommendations resultingfrom this study are:

1. Offset mechanical splices are not recommended for usewith bar sizes greater than No. 5 unless they can be shown tosatisfy the performance criteria;Fig. 9—Soffit of each specimen following testing.

487ACI Structural Journal/July-August 2008

2. Offset mechanical splices should not be used inapplications subject to seismic load reversals;

3. Offset mechanical splices should be included in the secondcategory of mechanical splices (having a fatigue limit of 83 MPa[12 ksi]) in AASHTO LRFD (2004), Table 5.5.3.4-1;

4. When used, offset mechanical splices must be designedto transfer 100% of the expected splice forces. They must notbe used to augment the capacity of traditional embedmentlaps splices;

5. The presence of confinement is expected to improvebehavior of these types of splices; a study of this effect isrequired. Modifications may need to be made to currentcodes to address the amount of confinement required to allow thesplice to properly function embedded in concrete; and

6. Coogler et al. (2006) have provided recommended revisionsto the Pennsylvania Department of Transportation specificationsand ASTM A1034-05b (ASTM International 2005) to allowthe inclusion of offset mechanical splices.

ACKNOWLEDGMENTSThe test program described in this paper was sponsored by the Pennsylvania

Department of Transportation through Contract #2006-033-040110. Thein-kind support of BarSplice Products Inc. and Erico International Inc. isgratefully acknowledged. The authors wish to acknowledge their discussionswith C. Paulson, Chair of ACI Committee 439, Steel Reinforcement; G. Weldonof Caltrans; and S. Holdsworth of BarSplice.

REFERENCESAASHTO ASD, 1996, “ASD Bridge Design Specifications,” seventeenth

edition, American Association of State Highway and Transportation Officials,Washington, DC.

AASHTO LFRD, 2004, “Bridge Design Specifications,” third edition,American Association of State Highway and Transportation Officials,Washington, DC, 1522 pp.

AASHTO M31, 1996, “Deformed and Plan Billet-Steel Bars forConcrete Reinforcement, American Association of State Highway andTransportation Officials, Washington, DC, 11 pp.

ACI Committee 318, 2005, “Building Code Requirements for StructuralConcrete (ACI 318-05) and Commentary (318R-05), American ConcreteInstitute, Farmington Hills, MI, 430 pp.

ACI Committee 439, 2007, “Types of Mechanical Splices for ReinforcingBars (ACI 439.3R-07),” American Concrete Institute, Farmington Hills,MI, 20 pp.

ASTM A1034-05b, 2005, “Standard Test Methods for TestingMechanical Splices for Steel Reinforcing Bars,” ASTM International,West Conshohocken, PA, 6 pp.

Caltrans, 2004, California Test 670 Method of Tests for Mechanical andWelded Reinforcing Steel Splices, California Department of Transportation,Sacramento, CA, 8 pp.

Coogler, K. L.; Harries, K. A.; and Gallick, M., 2006, “Evaluation ofOffset Mechanical Reinforcing Bar Splice Systems,” FHWA ReportNo. FHWA-PA-2006-033-040110, and University of Pittsburgh ReportNo. CE/ST-35, 91 pp.

CSA International, 2000, “Canadian Highway Bridge Design Code,”CAN/CSA-S6-00, Toronto, ON, Canada.

Harajli, M., 2007, “Cyclic Response of Concrete Members withBond-Damaged Zones Repaired using Concrete Confinement,” Materialsand Structures, V. 21, pp. 937-951.

Helgason, T., and Hanson, J. M., 1974, “Investigation of Design FactorsAffecting Fatigue Strength of Reinforcing Bars—Statistical Analysis,”Abeles Symposium on Fatigue of Concrete, SP-41 American ConcreteInstitute, Farmington Hills, MI, pp. 107-138.

Lukose, K.; Gergely, P.; and White, R. N., 1982, “Behavior of ReinforcedConcrete Lapped Spliced under Inelastic Cyclic Loading,” ACI JOURNAL,Proceedings V. 75, No. 5, July-Aug., pp. 355-365.

MacKay, B.; Schmidt, D.; and Rezansoff, T., 1988, “Effectiveness ofConcrete Confinement on Lap-Splice Perfomance in Concrete Beamsunder Reversed Inelastic Loading,” Canadian Journal of Civil Engineering,V. 16, No. 1, pp. 36-44.

Paulson, C., and Hanson, J. M., 1989, “A Summary and Review ofFatigue Data for Mechanical and Welded Splices in Reinforcing Bars,Structural Materials,” Proceedings of the Sessions Related to StructuralMaterials at Structures Congress’89, American Society of Civil Engineers,pp. 382-391.

Paulson, C., and Hanson, J. M., 1991, “Fatigue Behavior of Welded andMechanical Splices in Reinforcing Steel,” Final Report, Project 10-35,National Cooperative Highway Research Program.

Viwathanatepa, S.; Popov, E. P.; and Bertero, V. V., 1979, “Effects ofGeneralized Loadings on Bond of Reinforcing Bars Embedded in ConfinedConcrete Blocks,” UCB/EERC-79/22, Earthquake Engineering ResearchCenter, University of California-Berkeley, Berkeley, CA, 316 pp.

105-s46

Documents