Shear Behavior of Single Cast-in Anchors in Plastic Hinge Zones
International Journal of Concrete Structures and Materials
Shear Behavior of Single Cast-in Anchors in Plastic Hinge Zones
This paper presents two shear tests of -in. diameter cast-in anchors embedded in the plastic hinge zone of reinforced concrete columns. Design codes, such as ACI 318-14, require special reinforcement for concrete anchors in concrete that could be substantially damaged during an earthquake. The test anchors in this study were equipped with the anchor reinforcement recommended and verified in the literature. The column specimens were subjected to quasi-static cyclic loading before the test anchors were loaded in shear. Steel fracture was achieved in both test anchors despite cracks and concrete spalling occurred to the concrete within the plastic hinge zones. Meanwhile, the measured anchor capacities were smaller than the code-specified capacity, especially for the anchors subjected to cyclic shear. Concrete cover spalling was found critical to the observed capacity reduction, which caused combined bending and shear action in the anchor bolts. Measures should be developed to mitigate such adverse impact. In addition, further studies are needed for post-installed anchors before practical applications.
cast-in anchors; anchor reinforcement; concrete anchor; studs; fastening to concrete; reinforced concrete; seismic design
Concrete anchors are needed to connect structural steel
members and concrete. For example, reinforce concrete
frames have been strengthened using steel braces
and Jirsa 1990)
, in which, the steel braces can be fixed to
concrete beam/column ends using concrete anchors such as
post-installed adhesive anchors or through bolts. The anchor
bolts in such connections on the side faces of a concrete
frame are subjected to cyclic shear during an earthquake, and
the anchors have limited edge distance in the loading
Typical failure modes for such anchors in shear are anchor
steel fracture and concrete breakout failure (ACI 318, 2014).
Anchor steel failure is caused by fracture of an anchor shaft
in shear while concrete breakout failure is marked by a
concrete cone broken away from the base concrete, in which
the connection is located. Concrete breakout failure in shear
occurs when anchors are located close to an edge. Concrete
breakout is a brittle failure mode and thus not preferred for
anchor connections in seismic zones
(Petersen et al. 2013a)
The well-established design procedures for concrete anchors,
such as those stipulated in ACI 318-14, do not apply to the
anchors installed in plastic hinge zones. This is because the
concrete in plastic hinge zones likely develops substantial
damaged during an earthquake while the design procedures
are based on experimental tests of anchors in concrete that is
not significantly stressed/cracked.
Anchor reinforcement must be provided for the anchors
installed in plastic hinge zones. V-shaped hairpins encasing
the anchor shaft, as recommended in ACI 318-14 codes, may
not be practical for embedded connections. With a goal to
confine concrete around the shaft of an anchor in shear,
Petersen and Zhao (2013)
proposed new anchor shear
reinforcement, consisting of closely spaced stirrups, corner bars,
and crack-controlling bars. The proposed reinforcement has
been proved effective for anchors placed in
unstressed/uncracked concrete. Meanwhile, beam/column ends, where
anchor connections are placed, are likely to develop
significant damages during an earthquake
(Ibarra and Krawinkler
. The impact of potential severer damage in concrete
on the effectiveness of the anchor reinforcement and on the
behavior of anchors installed in plastic hinge zones has not
been evaluated. In this paper, laboratory tests were presented
to evaluate the recommended reinforcement for anchors
installed in the plastic hinge zone of concrete columns.
2. Literature on Anchor Behavior in Damaged Concrete
The behavior of concrete anchors as a connection between
steel members and concrete has been extensively studied as
Cook and Klingner (1992)
et al. (2005)
Eligehausen et al. (2006)
. Meanwhile, very limited studies are available
in the literature regarding the behavior of anchors in
significantly damaged concrete. A brief review of the studies of
anchors in cracked concrete is provided below followed by a
review of existing anchor reinforcement designs. Only
limited studies on post-installed anchors are included though
significant effort has been devoted to their capacities in crack
(Eligehausen et al. 2006)
2.1 Studies of Anchors in Damaged Concrete
Oehlers and Park (1992)
presented 25 push-off tests of
shear studs in a series of studies of shear transfer in
composite girders. At failure loads, concrete in front of the stud
crushed, leading to combined bending and shear in the studs.
The authors concluded that the longitudinal cracks near the
crush-prone region reduced the strength of concrete under
tri-axial compression. In this case, the stud capacity is likely
reduced by 10 percent. Transverse reinforcement near the
shear studs were found beneficial to the strength of the shear
studs because it confined the tri-axial compression zones.
The effect of steel reinforcing bars near studs was also
Saari et al. (2004)
in eight tests of shear studs
for use in composite construction. Closed stirrups were
provided to several specimens along with longitudinal bars
at all corners. Concrete failure occurred to the unreinforced
anchor (Specimen No. 2) while steel fracture occurred to the
reinforced anchor (Specimen No. 6) with capacity increase
about 100 percent. The ductility of the reinforced studs was
greatly improved, as compared with the unreinforced studs.
Zhang et al. (2001)
summarized the results of a
comprehensive study of anchors in cracked and uncracked concrete
led by Dr. Klingner at the University of Taxes, Austin in
1990s. The tests focused on expansion anchors and undercut
anchors that are of interest by nuclear industries. Groups of
four anchors were subjected to cyclic moments at a rate
similar to that in an earthquake. The tests indicated that the
cracks up to 0.25 mm [0.01 in.] did not affect the capacity of
tested undercut anchors but caused larger displacements at
failure. Meanwhile, the expansion anchors in cracks
developed low capacity and in some cases the failure modes
changed from concrete breakout (for anchors in uncracked
concrete) to pull-out (for anchors in cracked concrete).
2.2 Anchor Reinforcement Design
Anchor reinforcement must be provided for anchors in
concrete with severe damage. The methods for proportioning
anchor shear reinforcement are summarized in
. The currently recognized anchor shear
reinforcement includes hairpins and surface reinforcement, as
summarized in Fig. 1. Hairpins are deemed effective because
they can be placed close to the anchor shaft using a small
(Klingner et al. 1982; Lee et al. 2010)
transfer of shear load to surface reinforcement is usually
visualized using a strut-and-tie model (STM). Note that
strut-and-tie models usually permit the use of large size
reinforcing bars located at a large distance from the anchor
bolt. However, the tests by Nakashima (1998) have indicated
that reinforcing bars placed closer to the anchor are more
effective. As a result, the existing design codes such as ACI
318-14 require the anchor reinforcement to be within a
distance equal to half of the front edge distance (0.5ca1) as
illustrated in Fig. 1.
Petersen and Zhao (2013)
evaluated the behavior of cast-in
anchors with code-complying anchor reinforcement using
experimental tests. The authors proposed to use closed
stirrups in place of surface reinforcement. The new
reinforcement design, shown in Fig. 2, stresses the importance of
protecting concrete around the anchors using closely spaced
stirrups. With the concrete confined around the anchor, it is
expected that the concrete will provide shear resistance to
the anchor shaft.
Stirrups, as anchor shear reinforcement, are proportioned
using the anchor steel capacity in shear as specified by ACI
318-14, and the nominal yield strength of reinforcing steel.
Two stirrups are recommended next to the anchor shaft,
where the breakout crack in concrete may initiate under a
shear force. The rest of the required stirrups are placed with a
center-on-center spacing of 51 mm [2 in.] to 76 mm [3 in.]
within a distance of ca1 as shown in Fig. 2. A smaller
spacing may be used provided that the clear spacing
requirements are satisfied. The development length
requirements for the horizontal legs of the closed stirrups (as anchor
shear reinforcement) is satisfied through the interaction
between the stirrups and longitudinal bars at all four corners.
Nevertheless, a minimum length of 8db (db is the diameter of
reinforcing bars) was recommended for the horizontal legs
of the closed stirrups on both sides of the anchor as shown in
Crack-controlling bars include those at all four corners of
the closed stirrups as well as other bars distributed along the
concrete surfaces. Crack-controlling bars can be
proportioned using strut-and-tie models as shown later in specimen
design. These bars also need to be fully developed at both
Fig. 1 Anchor reinforcement for anchors in shear in ACI
318-14. a Hairpin anchor reinforcement and b surface
reinforcement with edge reinforcement.
sides of the anchor bolt, and a 90-degree bend as shown in
dashed lines in Fig. 2 may be used.
In summary, the existing anchor studies in the literature do
not reflect the damage concrete may experience in plastic
hinge zones. In addition, anchor reinforcement has been
proposed and verified using the tests in undamaged concrete.
This paper presents two tests of cast-in anchors with the
recommended anchor reinforcement in substantially
3. Experimental Program
3.1 Reinforced Concrete Column Specimen
A total of three reinforced concrete (RC) columns were
used for the single anchors in plastic hinge zones. The RC
columns had a section of 305 9 305 mm [12 9 12 in.] and
a height about 1.5 m [5 ft.] from the top face of the base
block, as shown in Fig. 3. The column base block had a
dimension of 1219 9 240 mm [48 9 20 in] and a height of
432 mm [17 in.] Two tie-down holes were created using
embedded PVC tubes. The tie down points were 0.91 m
[3 ft.] from the center of the column, following the hole
pattern in the strong floor of the University of
WisconsinMilwaukee (UWM) Structures Laboratory. The horizontal
loads were applied to the top of the column at 1.57 m
[62 in.] from the base through a steel loading block.
Eight No. 5 bars (ASTM Grade 60) were provided as the
longitudinal reinforcement for the columns. The concrete
cover was 38 mm [1.5 in.], typical for RC members with
exterior exposure. A section analysis with the actual material
properties shown below indicated that the column had a
nominal moment capacity about 105.8 kN-m [78 k-ft.]. This
Fig. 3 Dimension and reinforcement of shear sepcimens (1
in. = 25.4 mm).
corresponded to a lateral load capacity about 67.4 kN
[15 kips] at the top of the column. A shear design calculation
for this ultimate load led to No. 4 ties at a spacing of
152 mm [6 in.], as illustrated in Fig. 3.
3.2 Test Anchors
Two anchors are placed in each column specimen and
loaded in shear simultaneously. The test anchors were made
from 19-mm [ -in.] diameter ASTM A193 Grade B7
threaded rods. The net shear area (Asa;V ) for the threaded
rods was 2.2 cm2 [0.334 in.2]. Each test anchor, if fully
developed, has an ultimate shear capacity (Vu) of 117 kN
[26.4 kips]. Two single anchors, one installed on each side of
the RC column, were loaded simultaneously to eliminate the
torsion to the column. The test anchor had an embedded
length (hef ) of 203 mm [8 in.], and located at the middle of
the column side faces. With this configuration, one anchor
would have a front edge distance of 138 mm [5.4 in.] while
the other 167 mm [6.6 in.], as shown in Fig. 3. This
difference in the front edge distance was expected to create
small behavioral difference because of the anchor shear
reinforcement placed near the test anchors.
The test anchor was installed 203 mm [8 in.] above the
base block. The length of the plastic hinge zone was
assumed as 0.46 m [18 in.] (that is 1.5 times the column
section height), as shown by shaded region in Fig. 3. The
anchor position was selected assuming a crack spacing of
51 mm [2 in.] such that a major flexural crack would pass
through the test anchor if the anchor reinforcement were not
The design of the anchor shear reinforcement using
strutand-tie models is illustrated in Fig. 4. The required anchor
reinforcement for the 19-mm [ -in.] anchors in shear was
found to be 2.8 cm2 [0.44 in.2], following the design method
described in the previous section. Two No. 4 ties (with two
stirrups legs perpendicular to the anchor as the load-carrying
reinforcement) were provided: one located 51 mm [2 in.]
above the test anchor and the other below the anchor, as
shown in Fig. 3. The stirrups were not placed next to the test
anchor as recommended by
Petersen and Zhao (2013)
a space limitation and the need for crack-controlling
reinforcement. Two additional stirrups, located about 127 mm
[5 in.] from the test anchors as shown in Fig. 3, may also be
counted as anchor shear reinforcement; however, they were
expected to be less effective compared with the two adjacent
Crack-controlling reinforcement is important for the load
carrying reinforcement to function. With the configuration
shown in Fig. 4, the vertical splitting force from each anchor
in shear, which might cause horizontal splitting cracks, was
calculated as about 27 kN [6 kips]. The required
crackcontrolling reinforcement was 1.1 cm2 [0.17 in.2] using
0:6fy where fy is the yield strength of the bars. One No. 4
hairpin in the vertical plane, as shown in Fig. 3, was
provided. Another No. 4 hairpin in the vertical plane was
provided to resist the splitting force from the other anchor in
shear. In addition, the column was expected to develop
flexural cracks under lateral loading, especially in the plastic
hinge zones. The cracks might propagate passing the test
anchor. This cracking potential is in addition to that caused
by the force applied to the test anchors. One additional
hairpin was placed in the middle to restrain the flexural
cracking near the anchors. The development length
requirement for both the No. 4 ties and the U-shaped
hairpins was assumed satisfied through the interaction between
the closed ties and corner bars.
Fig. 4 Strut-and-tie models for the specimen design (1
in. = 25.4 mm; 1 kip = 4.45 kN).
Ready mixed concrete with Wisconsin Department of
Transportation Type A-FA mixture was used. The specified
concrete compressive strength was 27.6 MPa [4000 psi].
The hardened concrete had a compressive strength of
40 MPa [5800 psi] at 28 days, and the compressive strength
went up to 47.6 MPa [6900 psi] at about 84 days, when the
tests presented in this paper were conducted.
Eight No. 5 reinforcing steel bars with a nominal yield
strength of 414 MPa [60 ksi] were used as longitudinal
reinforcement. The measured stress–strain relationship,
shown in dashed lines in Fig. 5a experienced a grip slip. A
correction was performed based on the strain gage readings
and the strains calculated from the measured elongation over
a 203-mm [8-in.] gage length. The corrected stress–strain
relationship, shown in solid lines in Fig. 5a, indicates a yield
strength of 448 MPa [65 ksi] using the 0.2% offset method.
The ultimate strength of the reinforcing bars was measured
696 MPa [101 ksi] at a strain about 0.15.
The stress–strain relationship of ASTM A193 Grade B7
rods was measured using the tensile test of a coupon made
from a rod. Again, a grip slip was observed when the stress
was above 620 MPa [90 ksi], as shown in Fig. 5b. No strain
gage was used in this test; hence a correction was not
performed. The measured yield strength was 724 MPa [105 ksi]
corresponding to a 0.2 percent residual strain. The measured
ultimate strength was 914 MPa [132 ksi] at a strain about
3.4 Test Setup and Loading Protocol
The test setup is shown in Fig. 6. The column specimen
was fixed to the strong floor of the laboratory using two
high-strength threaded rods. An MTS Model 244.31,
245-kN [55-kip] actuator was used to apply a reversed cyclic
displacement at the top of the column. The lateral load was
applied to the column at 1.58 m [62 in.] above the base
block. Another MTS Model 244.31 actuator was used to
apply shear forces to the anchors through a loading adapter.
The centerline of this actuator was located at the height of
the test anchor. Strain gages and linear potentiometers were
used to monitor the behavior of the test columns and the
anchors. Specifically, the displacement at the column top
was measured using a linear potentiometer mounted on a
reference column. In addition, the anchor displacements
were measured using two linear variable differential
transformers (LVDT’s) mounted on the loading plate with respect
to the column surface as shown in Fig. 6. The
instrumentation details can be found elsewhere
(Petersen et al. 2013b)
The columns were subjected to reversed cyclic
displacements. The displacement history included groups of three
cycles with the peak displacements corresponding
to ± Dy, ± 2Dy, ± 3Dy, ± 4Dy, ± 6Dy, and ± 8Dy. The
yield displacement (Dy) was determined using a fiber-based
analysis of the column. The initial test results showed
differences between the command displacements and the actual
column displacements, and the peak displacements were
slightly modified accordingly. The loading rate for
Fig. 5 Materials used in test: a No. 5 Grade 60 rebars and
b ASTM A193 Grade B7 rods.
displacement cycles were kept at 6 mm/min [0.24 in./min]
throughout the tests.
The anchors in Specimen S1 were loaded in monotonic
shear till failure after the column was subjected to the
predefined loading history described above. The anchors in
Specimen S2 were loaded in reversed cyclic loading, and the
peak displacements were determined based upon the results
of the Specimen S1. The loading on the test anchors was in
the same direction as the load on the column, as illustrated
later in Fig. 9.
The test specimen could fail in several possible failure
modes under the complex combination of column loading
and anchor loading. Design checks for the column were
conducted for a variety of failure modes to ensure that
column specimen would not fail before the testing anchor
reached at its full design capacity. These failure modes
included flexural and shear failure of the column under the
load from the test anchors. In addition, shear friction along
flexural cracks was found unlikely to control the failure of
the test. Details about the design checks can be found
while similar design calculations are
recommended for design practices.
3.5 Shear Capacity Predication
The full shear capacity of the anchors, according to ACI
318-14, was calculated from
Vse ¼ 0:6futaAse;V
where futa is the ultimate tensile strength of the anchor steel
and Ase;V is the net shear area of the anchor shaft. However,
this calculation overestimated the actual shear capacity
because the cover concrete was lost during the test, and the
anchors had an exposed length. In this case, several
equations have been proposed to calculate the shear
capacity of anchor bolts with an exposed length l. For
Eligehausen et al. (2006)
proposed the following
while a different model was proposed by
Lin et al. (2011)
Vse ¼ 2
Vse ¼ fyaAse;v sinðbÞ þ
0:9Ase;v þ 3:4S
where fya is the yield strength of anchor steel, S is the section
modulus of the test anchor, and the rotation angle of the
exposed anchor (b) is estimated as h þ lp tan 1 emax, in which
the maximum tensile strain (emaxÞ was obtained from the
anchor material; h is the initial end rotation allowed by the
oversized holes and/or concrete deformation, lp is the length of
plastic hinge developed in anchor shaft at the ultimate load,
and may be taken as da (nominal diameter of the anchor).
To avoid friction between load plate and concrete surface
during the test, a 3-mm [1/8-in.] gap was controlled between
the loading plate and the concrete surface. Therefore, the
exposed length for the anchor in this study could be between
3 mm [1/8 in.] and 41 mm [1.6 in.], which was the concrete
cover depth. The calculated anchor shear capacity using Eq. 3
was thus between 64 kN [14.4 kips] and 122 kN [27.5 kips]
depending upon the damage to the concrete cover around the
test anchors. Equation (2) was proposed for anchors with a
minimum exposed length of d2a, where da is the anchor
diameter. Hence the predicted shear capacity was 41 kN [9.2 kips]
for the test anchors with an exposed length of 41 mm [1.6 in.].
4. Discussion of Test Results
4.1 Column Surface Cracks
The column in Specimen S1 had a typical hysteretic
behavior found in flexural members. An examination of the
strains in the middle longitudinal bar indicated that the first
yield occurred during the loading cycle at about 13 mm
[0.5 in.]. The largest column displacement was about
102 mm [4 in.], corresponding to 8 Deltay at a peak load
about 67 kN [15 kips]. The load–displacement behavior of
Column S2 was affected by the simultaneous loading to the
The crack map of Column S1 is shown in Fig. 7 after all
column loading cycles completed. While the crack maps
after each loading group can be found elsewhere
(Lin et al.
, a brief description of crack development is provided
as follows: flexural cracks first developed at the base and
within the plastic hinge zone. Smaller cracks were then
observed at higher location with the increase of the loading
level. The cracks in the plastic hinge zone widened and
propagated upon further loading. Note that the column base
crack is a result of yield penetration of steel reinforcing bars
in the foundation. The crack width at the base of the columns
was 0.3–0.5 mm [0.01–0.02 in.] at first yielding, which is
close to the bar yield slip predicted by
Zhao and Sritharan
. More cracks developed throughout the column
during the following loading cycles. The crack widths in
Fig. 7 were recorded after the 8 Deltay loading cycles. The
largest crack at the column base was 8.0 mm [0.31 in.],
which is slightly less than the ultimate bar slip recommended
Zhao and Sritharan (2007)
. Another crack at 102 mm
[4 in.] above the base had a width of about 7 mm [0.27 in.].
The strains in the middle longitudinal bar are shown in
Fig. 7a at the first peak of the 8Dy loading cycles, when
the crack widths in Fig. 7b were recorded. The crack widths
in millimeters, measured on the column faces, were listed in
Fig. 7 Strain in a longitudinal bar and crack map of Specimen
S1 column at 8Dy (Crack width in mm; 1
in. = 25.4 mm). a Strains in longitudinal bars and
b crack map (crack widths shown on the side of
Fig. 7b on the corresponding side of the column. The plastic
high zone, shown by the shaded region, was 610 mm
[24 in.] high, which is roughly twice the column section
height. The largest crack width was 0.4 mm [0.06 in.]
outside this region. Note that the design equations for concrete
anchors in ACI 318-14 were established based on tests of
both cast-in anchors and post-installed anchors in cracked
and un-cracked concrete; meanwhile, the crack width in the
tests have been limited to 0.5 mm [0.02 in.] to represent the
condition of concrete at serviceability limit state.
and Eligehausen 2008; Mahrenholtz et al. 2017)
observations in Fig. 7 support the current design regulations
that the equations in ACI 318-14 should be used outside the
plastic hinge zones, defined as twice the member depth from
a column or beam face.
4.2 Anchor Behavior in Shear
4.2.1 Monotonic behavior in Specimen S1
The behavior of the two single anchors subjected to
monotonic shear is shown in Fig. 8. The east anchor had a
slightly larger shear displacement as shown in Fig. 8a partly
because of a smaller edge distance. The X-axis thus shows
the average shear displacement measured relative to the
concrete front face in the load–displacement curve in
Fig. 8b. The Y-axis shows the total shear resistance from the
two single anchors. The anchors behaved elastically till
about 33.4 kN [7.5 kips], beyond which the slope reduced
significantly, indicating the crush of concrete in front of the
anchor shafts. Cracks on the east side face first showed up at
a load of 89 kN [20 kips] followed by spalling of concrete
cover as shown later in Fig. 10. Meanwhile, the cover cracks
on the west side were arrested by an existing flexural crack,
and the cracked concrete pieces were not significantly
crushed. This indicated that the west side anchor had a
relatively better lateral support than the east anchor, thus a
higher shear capacity.
The shear force increased with a relatively high slope
beyond 111.2 kN [25 kips] corresponding to a displacement
about 19 mm [0.75 in.]. Similar observation has been made
in a previous study
(Petersen and Zhao 2013)
: the anchor
shafts bent when the concrete cover spalled, and carried the
shear load partially in tension. Therefore, the final anchor
shear capacity had contributions from anchor shaft in shear,
bending, and tension as indicated by Eq. 3. The east anchor
fractured while the west anchor was also significantly bent,
when the combined load was 175 kN [39.4 kips]. On
average, each anchor had a shear capacity of 87.5 kN
[19.7 kips], which is within the predicted range shown
above. An after-test examination indicated that the
unsupported length of the anchor shafts was 38 mm [1.5 in.] at the
east side and 25 mm [1.0 in.] at the west side, which were
measured after the crushed concrete around the anchor shafts
The average shear capacity of 87.5 kN [19.7 kips] is much
higher than the code-specified breakout capacity (38 kN
[8.6 kips]), which would have controlled the anchor
behavior if the anchor reinforcement were not provided.
Anchor steel fracture was achieved because of the
wellFig. 8 Behavior of single anchor in plastic hinge zone under
monotonic shear loading. a Comparison of
measurments from two LVDT’s and b load vs average
confined core concrete in the plastic hinge zone of the
column specimen. The core concrete was protected with the
provided anchor reinforcement though the column outside
the concrete core had significant damage.
4.2.2 Cyclic behavior in Specimen S2
The anchors in Specimen S1 failed in shaft fracture after
the concrete cover spalled. The spalling was partially
attributed to the damage to the concrete in the plastic hinge
zone after the column was subjected to large inelastic
deformation. Anchor connections are loaded along with
concrete structures in reality; therefore the test of Specimen
S2 was slightly modified: both the column and the anchors
were subjected to loading simultaneously. The loading
protocol for the column stayed the same as Specimen S1, while
a displacement loading protocol was created based on the
observed behavior in Specimen S1 (Fig. 8b). Note that the
actuator that applied shear loading to the test anchors was
not controlled based on the measured anchor displacement
due to equipment limitations; hence the actual anchor
displacements were difference from the actuator displacement.
The cyclic behavior of Specimen S2 is plotted in Fig. 9a.
Again, the displacement in the plot is the average
Fig. 9 Behavior of single anchor in plastic hinge zone under
cyclic shear loading. a load vs anchor displacement
measured by LVDT’s and b load vs anchor
displacement measured by actuator.
displacement measured from two linear variable differential
transformers (LVDT’s), and the load is the combined shear
resistance from two test anchors. The overall anchor
behavior displayed pinching effects because the concrete,
crushed in front of the anchor shafts during the initial
loading, did not recover its lateral support to the shafts upon
unloading and reloading. In addition, the anchor behavior
was slightly under the shear loading in two directions: the
measured anchor shear capacity was 131 kN [29.4 kips]
when the actuators were in tension while the capacity was
157 kN [35.2 kips] when the actuators were in compression.
The lateral deformation of the column at 8 in. above base
and the related rotation may have been partly responsible to
the slight difference.
Unlike Specimen S1, the concrete cover spalling did not
occur at early stage of loading; hence the load–displacement
curve of Specimen S2 does not show an apparent ‘‘yield
point’’ at about 89 kN [20 kips], beyond which the loading
capacity remained unchanged with an increase in the anchor
displacement. The stiffness of Specimen S2 reduced
significantly at a higher load about 129 kN [29 kips]. Beyond the
point, the anchor behavior was stable for several groups of
displacement cycles from the load vs actuator displacement
curve in Fig. 9b. The actuator displacement contained the
lateral deformation of the column at 8 in. above the base.
With combined column-top loading and the anchor loading,
the column developed larger inelastic deformation than that
predicted by the analysis; hence the peak anchor
displacements stayed roughly the same during three groups of cyclic
loading, as shown in Fig. 9a.
Compared to the peak load of 175 kN [39.4 kips] in
monotonic test when the east anchor fractured, the anchor
rod on the west side in Specimen S2 fractured under a
loading of 157 kN [35.2 kips] in the opposite direction. The
fractured anchor in both tests had the smaller edge distance,
that is 137 mm [5.4 in.], as shown in Fig. 3. The smaller
ultimate load is attributed to the fact that the concrete
column in Specimen S2 experienced a smaller loading, 6Dy,
when the west anchor fractured. Consequently, the concrete
around the test anchors experienced less damage, thus
providing more lateral support to the anchor shafts compared
with Specimen S1.
4.2.3 Behavior of Confined Core
Concrete within the plastic hinge zones of Column S1 was
extensively cracked as shown in Fig. 10a before the anchor
was subjected to shear loading. Specifically, two major
cracks can be observed near the test anchors, as shown in
Fig. 7. These flexural cracks might have impacted the shear
behavior of the anchor if no crack-controlling were used.
The largest tensile strain in the U-shaped hairpins, measured
using a strain gage installed on the middle hairpin, was about
840 microstrains. Both the splitting force (Fig. 4) from
loading on the anchors and the flexural deformation of the
column contributed to the strain. This maximum strain
indicates that the design of the crack-controlling
reinforcement using 0:6fy is reasonable. With the crack-controlling
reinforcement, the large surface cracks did not extend
through the core concrete, as shown in Fig. 10b.
Wellconfined core concrete ensured the load transfer from the
anchor shafts to the concrete.
Concrete cover spalled in front of the anchor shaft, as
shown in Fig. 10. The spalling can be affected by the
inelastic deformation in the plastic hinge zones. With cover
spalling, as illustrated in the proposed anchor shear
reinforcement in Fig. 4, the shear behavior of concrete anchors
would be different from that in pure shear as assumed by
ACI 318 (2011). The double shear tests of anchor rods by
Lin et al. (2013)
provides understanding of the complex
stresses in an exposed anchor shaft. However, the model
failed accurately predicting the shear capacity of the anchors
in this study, likely because the exiting model does not
consider partial support from concrete cover before spalling.
The well confined concrete core with the proposed anchor
reinforcement allowed the anchors to develop improved
shear behavior, even though concrete in plastic hinge zone
experienced substantial damage. However, the observed
anchor shear behavior may not be desirable for the seismic
performance of diagonal braces because the full anchor
capacity would be achieved at a very large displacement.
The proposed anchor shear reinforcement protected well the
core concrete while further research is needed to protect
Design codes, such as ACI 318-14, requires special
reinforcement if anchors must be installed in concrete that can be
significantly damaged. This paper presents two shear tests of
cast-in anchors installed in the plastic hinge zone of concrete
column specimens. The anchor specimens in this study were
equipped with the recommended anchor reinforcement by
the authors, which emphasizes the importance of confining
core concrete and restraining concrete from cracking in
addition to distributing loads from the anchor to the rest of
the structure/structural element.
The anchors developed steel fracture instead of concrete
breakout failure despite large cracks and concrete spalling
occurred to the surrounding concrete. The successful tests
indicated that well confined core concrete, even within
plastic hinge zones, can support anchor connections in shear.
Meanwhile, concrete cover spalled during the tests,
adversely impacted the anchor shear behavior. Further
research is needed to protect cover concrete for connection
with concrete anchors to achieve desired behavior. In
addition, further studies are needed for post-installed anchors,
which are widely used in structural retrofit.
The study reported in this paper is from a project funded by
the National Science Foundation (NSF) under Grant No.
0724097. The authors gratefully acknowledge the support of
Dr. Joy Pauschke, who served as the program director for
this grant. Any opinions, findings, and recommendations or
conclusions expressed in this paper are those of the authors,
and do not necessarily reflect the views of NSF.
This article is distributed under the terms of the Creative
Commons Attribution 4.0 International License (http://
creativecommons.org/licenses/by/4.0/), which permits unre
stricted use, distribution, and reproduction in any medium,
provided you give appropriate credit to the original
author(s) and the source, provide a link to the Creative
Commons license, and indicate if changes were made.
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