Online ISSN 2093-6311
Print ISSN 1598-2351
International Journal of Steel Structures (2019) 19(1):319–328
https://doi.org/10.1007/s13296-018-0197-5
Force and Deformation Demands of Bolts in Steel Bolted Bracket
Moment Connections
Jong‑Kook Hong1
Received: 22 July 2018 / Accepted: 11 December 2018 / Published online: 14 December 2018
© Korean Society of Steel Construction 2018
Abstract
The behavior of pretentioned bolts at bracket-to-column in steel bolted bracket moment connections was investigated
through detailed 3D non-linear finite element analysis. Analysis results indicated the significant increase of axial force
in the bolt was caused by prying action due to column flange local bending. Tension force in the bolt was noted up to
166% of the initial bolt pretension, which corresponded to 125% of the nominal strength of the bolt. In addition, local
deformation at bolt threads resulted in high strain demand at nut-to-column flange, which is the potential hazard of
bolt fracture. A subsequent parametric study demonstrated that the strain concentration in the bolt could be shifted
away from the critical location to the unthreaded shank by simply adding a 25 mm thick gang washer on the back side
of column flange. But the tensile force demand in the bolts could not be reduced with additional gang washer plates.
Reducing initial bolt pretension (i.e., sung-tight condition) does not contribute to the reduction in bolt tension force
demand as well.
Keywords Bolted bracket · Finite element analysis · Pretention · Prying action · Steel moment connections
1 Introduction
Widespread damage, which was mainly associated with
brittle fracture in beam-to-column welded joints, was
observed during the 1994 Northridge, California earthquake due to the use of low toughness weld metals,
improper connection details, pre-Northridge construction practices, and so on. Because the connection failure
occurred at rotation levels well below the yield capacity
of the framing members, the perceptions regarding ductile
performance of steel moment frames were significantly
challenged. As a consequence, intensive researches were
conducted to develop connection details that ensure the
adequate seismic behavior and their products have demonstrated the successful performances.
Since the rigid frame action is the primary source of
resistance of lateral forces, a significant amount of research
focused on the welded steel moment resisting frames to form
the fully restrained (FR) connections and more stringent
* Jong‑Kook Hong
jkhong2002@gmail.com
1
R&D Center of Topinfra Co., Ltd, #613 Opulence Bldg., 254
Seocho‑daero, Seocho‑gu, Seoul 06647, Korea
requirements were imposed (AISC 2016b, c). Some of the
research programs paid their attention to an alternative
method by adopting high strength bolts and eliminating
the difficulties of field welding process (Gross et al. 2003;
Hantouche et al. 2013, 2015; Seek and Murry 2008). While
generally accepted connection details in seismic applications
consist of steel beams directly welded to steel columns at
their top and bottom flanges, bolted connections have been,
traditionally, considered as semi-rigid, and used to carry
the gravity loads. However, bolted connections have performed well during past earthquakes and the results of their
experimental tests supported that these connection types are
suitable for the seismic applications (Adan and Gibb 2009;
ICF Kaiser Engineers 2006; Sato et al. 2008; Sumner and
Murry 2002).
An approach that is readily applicable to repair as well
as new construction was attempted by Kasai et al. (1998).
The test results of a total of eight specimens, four W16 × 40
beam and W12 × 65 column subassemblages and four
W36 × 150 beam and W14 × 426 column subassemblages
with built-up hunch brackets, indicated that the bolted
bracket connections could be a viable alternative scheme
for repair of damaged welded moment connections. Both
experimental and analytical studies indicated that the
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maximum tensile force of the bolts in column was increased
by 30% due to prying action at bracket-column interface.
For conservatism, the design of bracket-to-column bolt
must be based on 1.3 times of the expected horizontal tensile force (Gross et al. 2003).
A total of seven full-scale moment connection specimens with cast iron brackets for either retrofit or new
construction scheme were conducted by Adan and Gibb
(2009). The beams were fabricated from W18 to W33
sections and the columns consisted of four W14 sections and three built-up box sections. All test specimens
demonstrated the satisfactory performance which met the
requirements of the AISC Seismic Provisions for Structural
Steel Buildings (AISC 2005). Based on this research, the
proprietary connection, Kasier Bolted Bracket moment
connections, were classified as the prequalified moment
connections in AISC 358: Prequalified Connections for
Special and Intermediate Steel Moment Frames for Seismic
Applications (AISC 2016c).
A series of full-scale tests were performed to evaluate
the cyclic performance of the rehabilitated moment connections at University of California, San Diego (UCSD) (Newell and Uang 2006). Among them, Specimen 3, which was
composed of two-sided W36 × 210 beams and a W27 × 281
column with a composite slab, used the bolted bracket rehabilitation scheme details with pre-Northridge top and bottom flange welding practice. On the first execution to 4%
drift, one of the bottom bolts fractured in brittle manner. All
bottom bracket bolts, then, progressively fractured before
the completion of two 4% drift cycles. Although this specimen met the AISC acceptance criteria for Special Moment
Frames (SMFs), the bolt fracture is not desirable in any
circumstance.
Steel moment connections with bolted brackets including the proprietary connections have demonstrated the reliable performance in the past experiments. However, one of
the experimental studies showed the unwanted non-ductile
bolt fracture. The objectives of this study are to investigate
the bolt performance at bolted bracket-to-column flange, to
identify the reason of bolt fracture, and to suggest an alternative to prevent the bolt fracture. For this study, the finite
element analysis program ABAQUS (ABAQUS 2014) was
used.
2 Finite Element Modeling
2.1 Representative Test Specimens
Figure 1 shows three representative test specimens having
bolted bracket moment connections with test configuration;
Specimen WH2 was one of eight test specimens conducted
by Kasai et al. (1998) and had single-sided beam with
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Fig. 1 Prototypes of bolted bracket test specimens
bottom bracket only. Specimen HH6 was one of seven test
specimens from ICF Kaiser Engineers (2006) and Adan and
Gibb (2009), and had single-sided beam with top and bottom brackets. No continuity plate was used in this specimen.
Specimen 3 was from Newell and Uang (2006) and had twosided beams with a total of four top and bottom brackets.
In addition, the composite slab was placed. Table 1 summarized the sectional properties of these specimens. Due to
high force demand from the larger beam sections, the largest
size of bolt (∅41 mm) was installed for the bracket-to-column connections in Specimen 3. Note that Specimen WH2
and Specimens HH6 showed the satisfactory performance
International Journal of Steel Structures (2019) 19(1):319–328
Table 1 Sectional properties of
bolted bracket specimens
Beam
Column
Mp ­ratioa
Bracket
Col. bolt
a
321
Specimen WH2
Specimen HH6
Specimen 3
W36 × 150 (A36)
[Mp = 2294 kN-m]
W14 × 426 (A572)
[Mp = 4899 kN-m]
2.14
Built-up
(A572)
∅38
(A490)
W30 × 108 (A572)
[Mp = 1956 kN-m]
W14 × 233 (A572)
[Mp = 2463 kN-m]
1.26
R020
(A148 Gr 80/50)
∅38
(A490)
W36 × 210 (A992)
[Mp = 4727 kN-m]
W27 × 281 (A992)
[Mp = 5279 kN-m]
1.12
R010C
(A148 Gr 90/60)
∅41
(A490)
The ratio of the flexural strength of column to the flexural strength of beam
while Specimen 3 failed in the connection bolts during the
execution to 4% story drift.
2.2 Modeling Scheme
Three-dimensional detailed finite element models were proposed to investigate the global and local behaviors of the
three representative bolted bracket test specimens. Models
were named after the specimen names (i.e., Model WH2
represented the reproduction of Specimen WH2). Model
geometry and boundary conditions reflected those used for
experimental testing. In Model HH6, the out-of-plane displacements at bracing points were constrained. For Model
3, the mid-width nodes of the beam top flange were constrained against displacement in out-of-plane direction at
every 300 mm, which was the shear stud spacing, to incorporate lateral stability provided by slab.
Due to large number of degrees of freedom and computing time involved in the models including detailed representations of bolted brackets, a simplified approach was
adopted. One of the bolted brackets per model was modeled
in detail using solid elements while the rest of the assembly
was modeled using 3D shell elements. Figure 2 shows an
example of the finite element meshes used in Model 3. The
solid bracket, bracket-to-column bolts, and a portion of the
adjacent beam and column were modeled with the combination of C3D8 eight-node hexahedral and C3D6 six-node
wedge elements. The general purpose S4R four-node shell
elements were used for the remainder of solid element beam
and column parts and S3 three-node shell element was used
for the shell bracket connections. Shell-to-solid coupling
technique was adopted at the transition from shell to solid
elements to enforce compatibility between different types
of finite elements. Table 2 summarizes the total number of
finite element with the element type used in the model.
The bolted bracket connection was not modeled explicitly. The bracket-to-beam bolts were not considered in
this study; instead, a tie constraint was applied between
bracket and beam bottom flange surfaces. The bolt hole in
the bracket-to-column was modeled 3.2 mm larger than the
Fig. 2 Finite element modeling
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Table 2 Finite element types and number of elements
Element type
Shell
Quadrilateral (S4R)
Triangular (S3)
Solid
Hexahedral (C3D8)
Wedge (C3D6)
Model WH2
Model HH6
3125
–
5770
122
44,950
15,174
33,712
8926
Model 3
6541
304
36,362
9846
Fu
Fy
E
y
3.5
y
Fig. 3 Stress–strain relationship
bolt diameter as it was. The hex bolt heads and nuts were
modeled as cylinders and the washers were not modeled.
The cylinder bolt shanks were considered without thread
part. All contact surfaces and anticipated contact surfaces
such as bolt head-to-bracket, nut-to-column flange, bolt
shank-to-bolt hole, and bracket-to-column were defined by
adopting a ‘master–slave’ type algorithm, which accounts
for friction and separation between interaction surfaces
(ABAQUS 2014). The slip coefficient of 0.30 was used to
account for Class A surface as defined in the AISC Specification for Structural Steel Buildings (AISC 2016a).
A tri-linear stress–strain relationship was utilized for
the models as shown in Fig. 3 (Mays 2000). An elastic
modulus, E, of 200 GPa and a Poisson’s ratio of 0.3 were
specified for the elastic material properties. The yield
strength, Fy, and the tensile strength, Fu, for each component in the model followed those obtained in the test. The
plasticity in the models was based on the von Mises yield
surface and associated flow rule. The plastic hardening
was defined by an isotropic hardening law.
The loading was applied in two steps; first, the pretention (659 kN for ∅38 mm bolts and 783 kN for ∅41 mm
bolts) was imposed simultaneously in all bolts based on
the minimum bolt pretension in the AISC Specification
for Structural Steel Buildings (AISC 2016a) to achieve
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clamping between the parts in the connection. In the second step, a monotonic loading up to 6% story drift was
applied at either column top or beam end just as in the test.
3 Connection Response
3.1 Global Response
Load versus displacements from analysis were compared
with their experimental data as shown in Fig. 4. The initial
elastic stiffness and the maximum load in the model corresponded well with the experimental result, although the
maximum load in Model WH2 was a bit higher than the
test data. The model simplification such as the tri-linear
stress–strain relationship, the uniform yield strength across
the entire section, one detailed bolted bracket model,
monotonic loading, actual versus model test configuration,
etc., contributed to these discrepancies. It was noted that
the strength degradation was observed in Models WH2 and
HH6, but not in Model 3.
Figure 5 shows the deformed shapes at 6% story drift.
Model WH2 showed top flange and web buckling in the
beam and this resulted in reduction of stress and strain
demands at beam flange-to-column CJP weld. Beam bottom flange deformation was restrained by maintaining the
right angle at the joint and shifted far away from the column face due to the bolted bracket. These results were
consistent with the experimental results in which the top
and bottom CJP welds survived during testing and the ductile fracture occurred at outermost net area of the beam
flange. In Model HH6, the beam top and bottom flange
deformations were noted at the tip of the brackets and the
column deformations were limited. These results also correlated well with the test results showing the plastic hinge
formed near the end of the bracket. Unlike Models WH2
and HH6 in which the column stiffness and strength were
relatively high compared to those of the beam, the column
deformation in Model 3 was remarkable and the separation
between the bracket and column flange was significant. But
the minimal deformation in the beam was noted. Considering the column kinking and bracket separation were also
observed in the test of Specimen 3, the overall responses
from the analysis reasonably represented those from the
experimental tests.
3.2 Bolt Tension Forces
Figure 6 shows the tension forces in the outermost bolt
at bracket-to-column connection at different story drift
stages. Dotted line representing the nominal strength of
the bolt, Rn, was added in the plot. The outermost row of
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Fig. 4 Global response comparison
bolts at each model showed similar behavior. The maximum bolt force in Model WH2 was 792 kN, which was
89% of the nominal bolt strength. But the maximum bolt
tension in Models HH6 and 3 was observed up to 955 kN
and 1300 kN, respectively, and these values corresponded
to 107% and 125% of the nominal strength of bolt. If compared with the initial pretensions (659 kN for ∅38 mm
bolts and 783 kN for ∅41 mm bolts), the bolt force was
120%, 145%, and 166% for Models WH2, HH6 and 3,
respectively. Results indicated that the bolt tension force
could be excessively increased and the most significant
increase was observed in Model 3.
3.3 Bracket‑to‑Column Flange Separation
The relative displacement between the bracket and the
column flange was monitored at end of vertical leg of
the bracket (see Fig. 7). It was noted that the separation
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Fig. 5 Deformed shape at story
drift of 0.06 rad
was small and almost constant even beyond 2% and
3% story drift in Models WH2 and HH6, respectively.
This is because the plastic hinge formation in the beam
concentrated the deformation demand to the beam and
released the deformation demand at other parts such as
column, bolt, and bracket. But the separation was continuously increased in Model 3. The beam deformation was
minimal; instead, the column deformation and bracket
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separation contributed to the rotational demand of the
system as also noted in Fig. 5c.
3.4 Prying Action in Column Flange
Kennedy et al. (1981) suggested the flange behavior models
based on the split-tee analogy. A flange was bolted to a rigid
support and a tension load was applied through a web. At the
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Separation (mm)
4
3
Model WH2
Model HH6
Model 3
2
1
0
0.0
0.01
0.02
0.03
0.04
0.05
0.06
Story Drift (rad)
Fig. 7 Bracket-to-column flange separation
Fig. 8 Thin-plate behavior in column flange
Fig. 6 Tension forces in bolt
low level of load, the flange behavior was somewhat rigid
and was so called “thick plate behavior”. Then, the sum of
the bolt tension force was equal to the applied load by statics. On the other hand, as the applied load became higher,
the plastic hinges were progressively formed in the flanges;
as a result, the bolts took an additional force due to prying
action as well as the applied load. It was called “thin plate
behavior”.
Column flange deformation in Model 3 as shown in Fig. 8
explains why the bolt force was significantly increased. The
deformation was compared at the centerline of outermost
row of bolts. The bracket and column flange were completely
clamped and no deformation could be observed just after
the pretension in the bolts applied. But, at 4% story drift,
the separation was apparent at bracket-to-column flange.
The separation was mainly caused by the column flange
local bending which was the indication of prying action.
The deformation in the bracket was negligible since the
bracket was relatively rigid. It is obvious that the “thin plate
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behavior” of the column flange was the major source of the
significant increase of bolt tension.
This phenomenon was not observed in Models WH2 and
HH6 where the beam was relatively flexible and week compared to the column. That’s because the beam yielded and
the plastic hinge formed in advance such that the strength of
the column never reached to the yield level. Throughout the
analytical simulation, the column flange connected through
bolts with the bracket showed as “thick plate behavior”.
4 Reduction in Bolt Demand
with Alternatives
4.1 Parametric Study
To reduce the high bolt demand observed in Model 3, a
parametric study was conducted. Three alternative models
were proposed by modifying Model 3; a 25 mm thick gang
washer was inserted between column flange and nut (back
side of column) in Model 3-1. Two 25 mm thick gang washers, one at column flange-to-nuts (back side of column) and
another at bracket-to-bolt heads (front side of bracket), were
added in Model 3-2. The initial pretension (= 783 kN/bolt)
in Model 3 was reduced to 223 kN/bolt in Model 3-Snug to
consider the snug-tight condition.
The yield strength, Fy, of the gang washer was assumed
to be 345 MPa. Contact surfaces due to the addition of gang
washers were also specified using the same methodology
adopted in Model 3.
4.2 Bolt Demand
4.2.1 Tension Force
Figure 9a shows the tension force comparison at the outermost bolt. At initial stage of loading till 2% story drift, the
bolt force in Models 3-1 and 3-2 increased slightly faster
compared to that in Model 3. But at larger drift stage, the
bolt force demands for all three models were practically
the same. The reduced bolt pretension in Model 3-Snug
increased even more rapidly and also converged into the
bolt force observed in Model 3 at 2% story drift and beyond.
It appears that the force demand in the bolt is not reduced
with this scheme. The flange local bending in column caused
the significant increase of tensile force in the connection
bolts. Either a change in column section with thicker flanges
or a stiffening the column flange would be effective in reducing the bolt tension force demand by minimizing prying
action.
Fig. 9 Bolt demands in alternatives
4.2.2 Deformation (Strain)
As noted in the experimental test of Specimen 3, the bolts
in bracket-to-column were progressively fractured at 4%
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story drift. After the test, it was reported that the bolt failure was initiated from local bending (kinking) in threaded
part of the bolt shank (Newell and Uang 2006). In the
analysis, as a quantitative measure of strain demand, the
plastic equivalent strain, PEEQ (ABAQUS 2014; El-Tawil
et al. 1999), was computed. PEEQ is defined as,
√
2
(1)
𝜀 𝜀
PEEQ =
3 ij ij
where εij is the plastic strain components in the directions
i and j.
As shown in Fig. 9b, the local nut deformation was
observed and the high PEEQ was indicated at the location of bolt threaded part in Model 3 and 3-Snug. This
is a potential hazard of bolt fracture since the strain and
stress concentrations are typically expected in elements
with discontinuity such as the bolt threads. When one or
two 25 mm thick gang washers as in Models 3-1 and 3-2
were added, the high strain region and the associated stress
concentration were moved away from the location of bolt
threads to the unthreaded shank. It is important to find that
shifting the bending deformation and the strain demand
away from the bolt threads could prevent the bolt fracture
observed in the test.
5 Conclusions
Steel bolted bracket moment connections have been suited
for the repair and new construction in high seismic region,
and the experimental tests have shown the satisfactory performance by providing rigid and ductile connections that
exceed the plastic rotation capacity shown in welded connections. A test conducted by Newell and Uang (2006),
however, reported the connection bolts fracture at the first
excursion to the 4% story drift although the connection
was able to meet the requirements specified in the AISC
Seismic Provisions for Structural Steel Buildings.
Detailed finite element analysis on three representative
test specimens were conducted using ABAQUS to study
the bolt behavior in the bolted bracket connections. From
the analytical study, the major findings are summarized
as follows:
1. Bolt tension force in bolted bracket-to-column flange
was noted up to 166% of the initial bolt pretension,
which corresponded to 125% of the nominal strength
of the bolt. Relatively week column flanges showed the
local bending in the connection and induced the separation between bracket and column flanges. Thin-plate
327
behavior, so called “prying action”, in the column flange
increased the bolt axial force demand significantly.
2. Bolt deformation in threaded part resulted in stress and
strain concentrations and could eventually be a potential
hazard of connection failure. Detailing preventing local
bending of the bolt in threaded part should be considered from the design stage to guarantee the satisfactory
performance.
3. Parametric study demonstrated that the bolt tension force
in bracket-to-column could not be reduced unless a column section with thicker flanges are used or adding stiffeners to the column flanges are considered. Reducing
initial bolt pretension (i.e., sung-tight condition) does
not contribute to the reduction of the tension force in
the connection bolts.
4. It is, however, believed that simply adding a gang
washer plate between column flange and nuts (back side
of column flange) such that the longer bolts are used
might be a solution since the bending deformation in
the bolts could be shifted away from the threads to the
unthreaded shank. Accordingly, the high strain demand
in the threaded part could be reduced.
Acknowledgements This study was made possible by the HDL Vibration and Sound Inc. The author would like to acknowledge Dr. HyunHun Choi and Dr. Young-Jong Moon for this constructive comments
on this study.
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