State-of-the-Art Review on Nonlinear Inelastic Analysis for Steel Structures

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Nov 15, 2013 (4 years and 5 months ago)

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State
-
of
-
the
-
Art Review on Nonlinear Inelastic
Analysis for Steel Structures

NRL Steel Lab., Sejong University

i

CONTENTS

1. INTRODUCTION

1

2. NONLIEAR INELASTIC ANALYSIS

3

2.1 Plastic
-
Zone Analysis

4

2.2 Quasi
-
Plastic Hinge Analysis

6

2.3 Elastic
-
Plastic Hinge Analysis

7

2.4 Notional
-

8

2.5 Refined
-
Plastic Hinge Analysis

9

3.

NONLINEAR INELASTIC EXPERIMENTS

11

3.1 Kanchanalai

s Two
-
Bay Frames

12

3.2 Yarimci

s Three
-
Story Frames

12

3.3 Avery and Mahendran

s Large
-
scale
testing

of Steel Frame Structures

13

3.4 Wakabayashi

s One
-
Quarter

Scaled Test of Portal Frames

13

3.5 Harrison

s Space Frame Test

14

3.5 Kim

s 3D Frame Test

14

4. DESIGN USING NONLIEAR INELASTIC ANALYSIS

15

4.1 Design Format

15

4.2 Modeling Consideration

16

4.2.1 Sections

16

4.2.2 Structural members

17

4.2.3 Geometric imperfection

ii

17

17

4.3 Design Consideration

18

-
carrying capacity

18

4.3.2 Resistance factor

19

4.3.3 Serviceability limit

19

4.3.4 Ductility requirement

20

REFERENCES

21

1.
INTRODUCTION

The steel design methods used in the U.S. are Allowable Stress Design (ASD), Plastic Design
(PD), and Load and Resistance Fac
tor Design (LRFD). In ASD, the stress computation is based on a
first
-
order elastic analysis, and the geometric nonlinear effects are implicitly accounted for in the
member design equations. In PD, a first
-
order plastic
-
hinge analysis is used in the stru
ctural analysis.
Plastic design allows inelastic force redistribution throughout the structural system. Since geometric
nonlinearity and gradual yielding effects are not accounted for in the analysis of plastic design, they
are approximated in member des
ign equations. In LRFD, a first
-
order elastic analysis with
amplification factors or a direct second
-
order elastic analysis is used to account for geometric
nonlinearity, and the ultimate strength of beam
-
column members is implicitly reflected in the desi
gn
interaction equations. All three design methods require separate member capacity checks including
the calculation of the K
-
factor.

This design approach is marked in Fig.

1

as the indirect analysis and

2

design method.

In the current AISC
-
LRFD Specifica
tion (AISC, 1994), first
-
order elastic analysis or second
-
order elastic analysis is used to analyze a structural system. In using first
-
order elastic analysis, the
first
-
order moment is amplified by B
1

and B
2

factors to account for second
-
order effects.
In the
Specification, the members are isolated from a structural system, and they are then designed by the
member strength curves and interaction equations as given by the Specifications, which implicitly
account for the effects of second
-
order, inelastici
ty, residual stresses, and geometric imperfections
(Chen and Lui, 1986). The column curve and beam curve were developed by a curve
-
fit to both
theoretical solutions and experimental data, while the beam
-
column interaction equations were
determined by a cu
rve
-
fit to the so
-
called "exact" plastic
-
zone solutions generated by Kanchanalai
(1977).

In order to account for the influence of a structural system on the strength of individual
members, the effective length factor is used as illustrated in Fig. 2.

The
effective length method generally provides a good design of framed structures.
However, several difficulties are associated with the use of the effective length method as follows:

(1) The effective length approach cannot accurately account for the intera
ction between the
structural system and its members. This is because the interaction in a large structural system is too
complex to be represented by the simple effective length factor K. As a result, this method cannot
accurately predict the actual requ
ired strengths of its framed members.

(2) The effective length method cannot capture the inelastic redistributions of internal forces in a
structural system, since the first
-
order elastic analysis with B
1

and B
2

factors accounts only for
second
-
order effe
cts but not the inelastic redistribution of internal forces. The effective length
method provides a conservative estimation of the ultimate load
-
carrying capacity of a large structural
system.

(3) The effective length method cannot predict the failure mo
des of a structural system subject to a
given load. This is because the LRFD interaction equation does not provide any information about

3

failure modes of a structural system at the factored loads.

(4) The effective length method is not user
-
friendly for

a computer
-
based design.

(5)

The effective length method requires a time
-
consuming process of separate member capacity
checks involving the calculation of K
-
factors.

With the development of computer technology, two aspects, the stability of separate members
,
and the stability of the structure as a whole, can be treated rigorously for the determination of the
maximum strength of the structures. This design approach is marked in Fig. 1

as the direct analysis
and design method

(Kim and Chen, 1996a
-
b)
. The dev
elopment of the direct approach to design is
called “Advanced Analysis” or more specifically, “Second
-
Order Inelastic Analysis for Frame
Design.” In this direct approach, there is no need to compute the effective length factor, since
separate member capac
ity checks encompassed by the specification equations are not required. With
the current available computing technology, it is feasible to employ nonlinear inelastic analysis
techniques for direct frame design. This method has been considered impractical

for design office
use in the past.

Over the past 20 years, extensive research has been made to develop and validate several
nonlinear inelastic analysis methods.

The purpose of this paper is to review recent efforts to develop
various nonlinear inelasti
c analyses ranging from a simple elastic
-
plastic to
rigorous

plastic
-
zone
analysis for frame design.

Emphasis in this review is
design application of nonlinear inelastic
analysis.

This paper
also summarizes

reports of experimental studies to provide inel
astic nonlinear
behavior of framed structures.

The analysis and design principle using nonlinear inelastic analysis

2. NONLINEAR INELASTIC ANALYSIS

4

Five different types of nonlinear inelastic analysis methods are discussed in t
he following:

(1) Plastic
-
zone method

(2) Quasi
-
plastic hinge method

(3) Elastic
-
plastic hinge method

(4) Notional
-
load plastic hinge method

(5) Refined
-
plastic hinge method

These different methods are based on the degree of refinement in representing th
e plastic
yielding effects. The plastic
-
zone method uses the greatest refinement while the elastic
-
plastic hinge
method allows a drastic simplification. The quasi
-
plastic hinge method is somewhere in between
these two methods. The notional
-
hinge method and the refined
-
plastic hinge method are
an improvement on the elastic
-
plastic hinge method for

approximating real behavior of structures
.
-
deformation characteristics of the plastic analysis methods are illustrated in Fig.
3
, while t
he
spread of plasticity is illustrated schematically in Fig.
4
.

2.1
Plastic
-
Zone Method

In the plastic
-
zone method, frame members are discretized into finite elements, and the cross
-
section of each finite element is subdivided into many fibers shown in Fig
.
5
. The deflection at each
division point along a member is obtained by numerical integration. The incremental load
-
deflection
response at each loading step, which updates the geometry, captures the second
-
order effects. The
residual stress in each fib
er is assumed constant since the fibers are small enough. The
s
tress
s
tate at
each fiber can be explicitly traced so the gradual spread of yielding can be captured. The plastic
-
zone
analysis eliminates the need for separate member capacity checks since it

explicitly accounts for
second
-
order effects, spread of plasticity, and residual stress. As a result, the plastic
-
zone solution is
known as an "exact solution." The AISC
-
LRFD beam
-
column equations were established in part
based upon a curve
-
fit to the "
exact" strength curves obtained from the plastic
-
zone analysis by
Kanchanalai (1977).

5

There are two types of plastic
-
zone analyses. The first involves the use of three
-
dimensional
finite shell elements in which the elastic constitutive matrix in the usual

incremental stress
-
strain
relations, is replaced by an elastic
-
plastic constitutive matrix when yielding is detected. Based on a
deformation theory of plasticity, the effects of combined normal and shear stresses may be accounted
for. This analysis requ
ires modeling of structures using a large number of finite three
-
dimensional
shell elements and numerical integration for the evaluation of the elastic
-
plastic stiffness matrix.
The three
-
-
of
-
plasticity analysis when combined with second
-
order theory which
deals with frame stability is computational intensive and, therefore, best suited for analyzing small
-
scale structures, or if the detailed solutions for member local instability and yielding behavior are
required. Since a detailed anal
ysis of local effects in realistic building frames is not common
practice in engineering design, this approach is considered too expensive for practical use.

The second approach for second
-
order plastic
-
zone analysis is based on the use of beam
-
column theo
ry, in which the member is discretized into line segments, and the cross
-
section of each
segment is subdivided into finite elements. Inelasticity is modeled considering normal stress only.
When the computed stress at the centroid of any fiber reaches the

uniaxial normal strength of the
material, the fiber is considered to have yielded. Also, compatibility is treated by assuming that full
continuity is retained throughout the volume of the structure in the same manner as elastic range
calculations. Altho
ugh quite sharp curvature may exist in the vicinity of inelastic portions of the
structure, “plastic hinges” can never develop. In plastic
-
zone analysis, the calculation of forces and
deformations in the structure after yielding requires an iterative tria
l
-
and
-
error process because of the
nonlinearity of the load
-
deformation response, and the change in cross
-
section effective stiffness in
inelastic regions associated with the increase in the applied loads and the change in structural
geometry. Although mo
st plastic
-
zone analysis methods have been developed for planar analyses
(Clarke et al., 1992; White, 1985; Vogel, 1985; El
-
Zanaty et al., 1980; Alvarez and Birnstiel, 1967)
three
-
dimensional plastic
-
zone techniques are also available (Wang, 1988; Chen and

Atsuta, 1977).

6

A plastic
-
zone analysis that includes the spread of plasticity, residual stresses, initial
geometric imperfections, and any other significant second
-
order effects, would eliminate the need for
checking individual member capacities in the
frame. Therefore, this type of method is classified as
nonlinear inelastic inelastic analysis in which the checking of beam
-
column interaction equations is
not required. In fact, the member interaction equations in modern limit
-
states specifications were

developed, in part, by curve
-
fit to results from this type of analysis. In reality, some significant
behaviors such as joint and connection’s performances tend to defy precise numerical and analytical
modeling. In such cases, a simpler method of analysi
s that adequately represents the significant
behavior would be sufficient for engineering application.

Whereas the plastic
-
zone solution is regarded as an "exact solution," the method may not be
used in daily engineering design, because it is too intensive

in computation. Its applications are
limited to (ECCS, 1984):

(1) The study of detailed structural behavior

(2) Verifying the accuracy of simplified methods

(3) Providing comparison with experimental results

(4) Deriving design methods or generating cha
rts for practical use

(5) Applying for special design problems

2. 2
Quasi
-
Plastic Hinge Method

The quasi
-
plastic hinge method developed by Attala (1994) is an intermediate approach
between the plastic
-
zone and the elastic
-
plastic hinge methods. It req
uires less computation but its
results are very similar to those of plastic
-
zone method. For this reason, it is called a quasi
-
plastic
hinge method.

An element, developed from equilibrium, kinematic, and constitutive relationships, accounts
astification under combined bending and axial force. Inelastic force
-
strain model of

7

the cross
-
section is developed by fitting nonlinear equations to data of the moment
-
axial force
-
curvature response. Using the inelastic cross
-
section model, flexibility
coefficients for the full
member are obtained by successive integrations along its length. An inelastic
-
element stiffness
matrix is obtained by the use of the incremental flexibility relationships.

Initial yield and full plastification surface are used to

analytically represent gradual yielding
effect of the cross
-
section. Ketter’s residual stress pattern (1955) is used to determine an initial yield
surface. Ketter’s pattern has peak compressive residual stresses at the flange tips equal to 0.3F
y

with
a
linear transition of stress from the flange tips to the web
-
joint and constant tensile stress through the
web. A fully plastic surface is generated by calibration to a plastic
-
zone solution (Sanz
-
Picon, 1992).
The parameters of the full plastification eq
uation are determined by a curve
-
fit procedure.

This method predicts strengths with an error less than 5% compared with the plastic
-
zone
method for a wide range of case studies. The accuracy of this method is thus compatible with the
plastic
-
zone method a
nd less computational effort is necessary.

However, it is difficult to extend this method to three
-
dimensional analysis since the
formulation is based on flexibility relationships.

As a result, it does not meet one of the
requirements of

nonlinear inelas
tic

analysis

of the SSRC task force report (1993), which states

The model should be readily extensible to three
-
dimensional analysis. That is, the framework of
the model should accommodate the formulation of three
-
dimensional elements.

Moreover, this
model does eliminate the necessity of the refined model through the cross
-
section but still requires
many elements along the member.

2. 3
Elastic
-
Plastic Hinge Method

A more simple and efficient approach for representing inelasticity in frames is the elas
tic
-
plastic hinge method. It assumes that the element remains elastic except at its ends where zero
-
length plastic hinges form. This method accounts for inelasticity but not the spread of yielding or

8

plasticity at sections nor the residual stress effect
between two plastic hinges.

The elastic
-
plastic hinge methods may be divided into; first
-
order and second
-
order plastic
analyses. For first
-
order elastic
-
plastic hinge analysis, the nonlinear geometric effects are neglected,
and not considered in the for
mulation of the equilibrium equations. As a result, the method predicts
the same ultimate load as conventional rigid
-
plastic analyses.

In second
-
order elastic
-
plastic hinge analysis, the deformed structural geometry is considered.
The simple way to acc
ount for the geometric nonlinearity is to use the stability function which enables
only one beam
-
column element per a member to capture the second
-
order effect. This provides an
efficient and economical method of frame analysis, and has a clear advantage
over the plastic
-
zone
method. This is particularly true for structures in which the axial force in component members is
small and the dominated behavior is bending. In such cases, second
-
order elastic
-
plastic hinge
analysis may be used to describe the in
elastic behavior sufficiently, assuming that lateral
-
torsional and
local buckling modes of failure are not prevented (Liew, 1992).

The second
-
order elastic
-
plastic hinge analysis is only an approximate method. When used
to analyze a single beam
-
column ele
ment subject to combined axial load and bending moment, it may
overestimate the strength and stiffness of the element in the inelastic range. Although elastic
-
plastic
hinge approaches provide essentially the same load
-
displacement predictions as plastic
-
z
one methods
for many frame problems, they may not be classified as
nonlinear inelastic

analysis methods in
general (Liew et al., 1994; Liew and Chen, 1991; White, 1993).

However, research by Ziemian (Ziemian et al., 1990; Ziemian, 1990) has shown that the
elastic
-
plastic hinge analysis can be classified as an advanced inelastic analysis since it is accurate for
matching the strength and load
-
displacement response of several building frames from plastic
-
zone
analysis. Many cases considered in Ziemian

s work, especially when the axial load is less than
0.5P
y
, are not sensitive benchmarks for determining the accuracy and the possible limitations of the
elastic
-
plastic hinge method. Therefore, suitable benchmark problems should be used to provide a

9

more

in
-
depth study of the qualities and limitations of second
-
order elastic
-
plastic hinge method
before it can be accepted as a legitimate tool in the design of steel structures.

For slender members whose dominant mode of failure is elastic instability, the m
ethod provides good
results when compared with plastic
-
zone solutions. However, for stocky members with significant
yielding, the plastic
-
hinge method over
-
predicts the actual strength and stiffness of members due to
the gradual stiffness reduction as the

spread of plasticity increases in an actual member (Liew and
Chen, 1991; Liew et al., 1991; White et al., 1991). As a result, considerable refinements must be
made before it can be used for analysis of a wide range of framed structures.

2. 4
Notional
-
Lo
-
Hinge Method

One approach to advance the use of second
-
order elastic
-
plastic hinge analysis for frame
design is to specify artificially large values of frame imperfections (i.e., initial out
-
of
-
plumbness
)
.
This is the approach adopted by EC3 (1
990) for frame design using second
-
order analysis. In
addition to accounting for the standard erection tolerance for out
-
of
-
plumbness, these artificial large
imperfections intend to account for the effect of residual stresses, frame imperfections, and dis
tributed
plasticity not considered in frame analysis. The geometric imperfections adopted by EC3 are a
maximum out
-
of
-
plumbness of
Ψ
0

= 1/200 for an unbraced frame, but no maximum out
-
of
-
straightness value recommended for a braced member as shown in Fig.
6
.

The notional load plastic hinge approach is similar in concept to the “enlarged” geometric
imperfection approach of the EC3. T
he ECCS (1984, 1991), the Canadian Standard (
1989,
1994),
and the Australian Standard (1990) allow to use this technique. The notional
-
equivalent lateral loads to approximate the effect of member imperfections and distributed plasticity
.
In the ECCS, the exaggerated notional loads of 0.5 % times gravity loads are used to avoid over
-
predicting the strength of the member as does the elastic
-
plastic hinge method. The application of
these notional loads to several example frames is illustr
ated in Fig.
7
. Liew'

s research (1992) shows

10

that this method under
-
predicts the strength by more than 20% in the various leaning column frames
and over
-
predicts the strength up to 10% in the isolated beam
-
columns subject to the axial forces and
bending
moments. As a result, modification of this approach is required before it may be used in
design applications.

2. 5
Refined Plastic
-
Hinge Method

In recent work by Abdel
-
Ghaffar et al. (1991), Al
-
Mashary and Chen (1991), King, et al.
(1991), Liew and Chen
(1991), Liew et al. (1993a
-
b), White et al. (1991), Kim (1996), Kim and Chen
(1996),
Chen and Kim (1997),
Kim and Chen (199
7
)
, Kim et al (2000) and
among others, an inelastic
analysis approach, based on simple refinements of the elastic
-
plastic hinge model
, has been proposed
for plane frame analysis. It represents the effect of distributed plasticity through the cross
-
section,
assuming that the plastic hinge stiffness degradation is smooth. The inelastic behavior of the
member is modeled in terms of membe
r force instead of the detailed level of stresses and strains as
used in the plastic
-
zone analysis model. The principal merits of the refined
-
plastic hinge model are
that it is as simple and efficient as the elastic
-
plastic hinge analysis approach, and it

is sufficiently
accurate for the assessment of strength and stability of a structural system and its component members.

The refined plastic
-
hinge method is based on simple modifications of the elastic
-
plastic hinge
method. Two modifications are made to a
ccount for the gradual section stiffness degradation at the
plastic hinge locations as well as gradual member stiffness degradation between the two plastic hinges.
Herein, the section stiffness degradation function is adopted to reflect the gradual yieldi
ng effect in
forming plastic hinges. Then, the tangent modulus concept is used to capture the residual stress
effect along the member between two plastic hinges. As a result, the refined plastic
-
hinge method
retains the efficiency and simplicity of the p
lastic hinge method without overestimating the strength
and stiffness of a member.

In the recent work by Liew (1992), the LRFD tangent modulus is used to account for both the

11

effect of residual stresses and geometric imperfections. This model does not acc
ount for geometric
imperfections when P/P
y

is less than 0.39, because the LRFD tangent modulus is identical to the
elastic modulus in this range. As a result, the approach over
-
predicts the column strength by more
than 5% when KL/r of the column is greate
r than 85 for yield stresses at 36 ksi, and when KL/r of the
column is greater than 70 for yield stresses at 50 ksi. The LFRD E
t

may not be an appropriate model
to be used for nonlinear inelastic analysis (Kim, 1996; Kim and Chen, 1996).

The CRC tangent
modulus in Liew's work (1992) only accounts for the effect of residual
stresses. It over
-
predicts the strength of members by about 20% compared to the conventional
LRFD solutions, because the modulus does not account for the effect of geometric imperfecti
ons.
However, in the CRC tangent modulus model, different members with different residual stresses can
be incorporated since the effect of geometric imperfections is considered separately. As a result,
CRC tangent modulus is used in refined plastic analy
ses.

Second
-
order inelastic analysis methods for the three
-
dimensional structure have been
developed by Orbison (1982), Prakash and Powell (1993), Liew and Tang (1998)
, Kim et al (2001),
Kim and Choi (2001) and Kim et al (2001)
. Orbison's method is an ela
stic
-
plastic hinge analysis
without considering shear deformations. The material nonlinearity is considered by the tangent
modulus
t
E

and the geometric nonlinearity is by a geometric stiffness matrix. Orbison's method,
however, und
erestimates the yielding strength up to 7% in stocky members subjected to axial force
only. DRAIN
-
3DX developed by Prakash and Powell is a modified version of plastic hinge methods.
The material nonlinearity is considered by the stress
-
strain relationship

of the fibers in a section. The
geometric nonlinearity caused by axial force is considered by the use of the geometric stiffness matrix,
but the nonlinearity caused by the interaction between the axial force and the bending moment is not
considered. This
method overestimates the strength and stiffness of the member subjected to
significant axial force. Liew and Tang's method is a refined plastic
hinge analysis. The effect of
residual stresses is taken into account in conventional beam
-
column finite element

modelling.

12

Nonlinear material behavior is taken into account by calibration of inelastic parameters describing the
yield and bounding surfaces. Liew and Tang's method, however, underestimates the yielding strength
up to 7% in stocky member subjected to ax
ial force only.

Against this background, it can be concluded that the refined
-
plastic hinge method strikes a
balance between the requirements for realistic representation of frame behavior and for ease of use.
It is considered that in both theses respects
, the method is satisfactory for general practical use.

3. NONLINEAR INELASTIC EXPERIMENTS

Experimental studies to capture inelastic nonlinear behavior of framed structures are
summarized. The frames riviewed herein were tested by Kanchanalai(1977), Y
arimci(1966),
Avery(1999), Wakabayashi(1972), Harrison(1964) and Kim and Kang(2001).

3.1 Kanchanalai

s Two
-
Bay Frames

Three two
-
bay full
-
size frames were tested to verify the Plastic
-
zone analysis(Kanchanalai,
1977). The dimensions and members of Frame 2

among these frames are shown in Fig. 8. The
material properties of the members are summerized in Table 1. The frames were designed to behave
equivalently to a one
-
story two
-
bay and could be tested on the floor. Supports were provided only at
the top and

bottom of the interior column member. All frames were bent with respect to the week
axis in order to avoid out
-
of
-
plane buckling. In Frame
2,

all columns were loaded simultaneously up
to about 70kips, corresponding to points 2
-
11 in Fig. 9. Then, only
the axial load on the interior
column was increased up to point 17, where the frame reached its instability limit load of 233.6 kips.
Comparisons of the test results with the plastic zone theory are shown in Fig. 9. In
general
, good
agreements are observe
d.

13

3.2 Yarimci

s Three
-
Story Frames

An experimental research study was conducted at Lehigh University for three full
-
size frames
(Yarimci, 1966). Fig. 10 shows dimensions and loads conditions of Frame C among the three frames.
To investigate and compare

the mechanical properties of the members with nominal values, Yarimci
conducted a series of seven beam tests. The results of these tests are summarized in Table 2. The
beams were welded to the columns and designed so as to behave elastically in the wors
condition: the flexibility of the connections was eliminated from a factor which affects the strength of
the frames. The frames were sandwiched and supported laterally by two parallel auxiliary frames
preventing out
-
of
-
plane buckling. All membe
rs were bent in strong axis. The result of test is
shown in Fig. 11 for Frame C. The load deflection behavior at the first and third story is shown in
Fig. 11.

3.3 Avery and Mahendran

s Large
-
Scale Testing
of

Steel Frame Structures

A series of four te
sts was conducted by Avery and Mahendran(1999). Each of the four
frames could be classified as a two
-
dimensional, single
-
bay, single
-
story, large
-
scale sway frame with
full lateral restraint and rigid joints, as shown in Fig. 12.
In Frame 2, Non
-
compact
I
-
sections(310UB32.0) of Grade 300 steel(nominal yield

stress=320MPa)was used.

This section was
selected as one of the standard hot
-
rolled I
-
sections mostly affected by local buckling. The
dimensions, material properties, and section properties used in F
rames 2 are listed in Table 3. The
vertical and horizontal loads were applied simultaneously in a ratio of approximately four times
greater than the horizontal reaction measured by the load cell. The frame failed by in
-
plane instability
due to a reduced s
tiffness caused by yielding and P
-
Δ

effect. The horizontal reaction force and the
measured relative in
-
plane horizontal displacement of the right hand column for test Frame 2 are
related in Fig. 13.

14

3.4 Wakabayashi

s One
-
Quarter Scaled Test of Portal Fra
mes

Two
-
series of test were conducted for a one
-
story frame and a two
-
story frame by
Wakabayashi et al(1972). Configurations of the two
-
story frame are shown in Fig. 14. The
nominal dimensions of members are H
-
100

100

6

8 for columns and H
-
100

50

4

6 fo
r beams.
The specimens consist of rolled H
-
shapes. The connections were welded and stiffened to prevent
local buckling in the joint panels. To prevent the out
-
of
-
plane buckling, two of the same specimens
were set in parallel and connected at the joints
and the mid
-
length of the members. In the other
words, twin specimens were tested simultaneously. Measured Material and sectional properties of
members are listed in Table 4.

The vertical load was first applied at the top of four columns by a fixed tes
ting machine.
The
parallel

twin specimens were loaded simultaneously. Then, the horizontal load at the top of
frame was increased gradually. When the frame swayed by the horizontal loading jack followed a
horizontal movement so that vertical loading poi
nts could be kept on the center of the columns. The
were measured by the load cells which were installed between the hydraulic jacks and the
specimen.

-
deflection curves of the two
-
story frames are shown in Fig. 15. Comparisons of a
series

of test show the effects of axial force and stiffness of the beam on the frame behavior. The
larger the axial force in columns and the smaller the stiffness of the beam, the more unstable the
frames become.

3.5 Harrison

s Space Frame Test

The equilater
al triangular space frame depicted in Fig. 3 was tested by Harrison(1964)
in the

J.W.
Roderick Laboratory for Materials and Structures at the University of Sydney. Configuration of the
frame is shown in Fig. 16. Measured dimensions and

material properties

are listed in Table 5. A

15

horizontal load(H) is applied on the top of the column and a vertical load of 1.3H is applied at mid
span of the beam.

It can be seen from Fig. 17 that, compared to the experimental results, the plastic
-
zone
analysis predicted a

slightly stiffer response of the space frame under the applied loads. As the
column bases of the space frame were welded to steel plates clamped to steel joists(Harrison 1964),
the more flexible response measured in the laboratory test might have been ca
used by the flexibility of
the joist flanges.

3.6 Kim

s 3D Frame Test

Two
-
series of test were conducted for space steel frame subjected proportional loads shown
in Fig 18 and space steel frame subjected proportional loads shown in Fig.

19 by Kim and Ka
ng(2001).
Hot
-
rolled I
-
section was used for all three frames. Nominal dimension of
the section was H
-
150
×
150
×
7
×
10 commonly used in Korea. The dimensions and properties of the
section
are listed in Table 6. The section is compact so that it is not suscept
ible to local buckling.

For proportional loads test, The vertical loads were applied on the top of the four columns,
and the horizontal
loads were applied on the column

and

at the second floor level of the test
frame. The vertical loads were slowly increased until the system could not resist any more loads.
The horizontal loads were automatically increased according to the specified l
oad ratio for each test
frame controlled by the computer system.

For non
-
proportional loads test, The vertical loads were applied on the top of the four
columns, and the horizontal load was applied on the column

at the second floor level of the test
fr
ame. The vertical loads were first increased 680
kN

and maintained during the experiment. The
horizontal load was slowly increased until the test frame could not resist any more loads.

Fig. 20. and Fig. 21. show load
-
displacement cu
rve for test frames. The obtained results
from 3D non
-
linear analysis and AISC
-
LRFD method were compared with experimental data.
ABAQUS, one of mostly widely used and accepted commercial finite element analysis
program
, was

16

used.
s obtained by the experiment and AISC
-
LRFD method are compared
in Table 7 and 8. The results showed that the AISC
-
LRFD capacities were approximately 25 percent
conservative for frame subjected to
proportional

loads test and 28 percent conservative for non
-
proportional

loads test. This difference is
derived

from the fact that the AISC
-
LRFD approach does
not consider the inelastic moment redistribution, but the experiment includes the inelastic
redistribution effect.

4. DESIGN USING NONLINEAR INELASTIC AN
ALYSIS

4
.1 Design Format

Nonlinear inelastic analysis follows the format of Load and Resistance Factor Design. In
A
ISC
-
LRFD
(1994)
, the factored load effect does not exceed the factored nominal resistance of
structure. Two kinds of factors are used:
one is applied to loads, the other to resistances. The load
and resistance factor design has the format

i i n
Q R
  

(1)

where
n
R

= nominal resistanc
e of the structural member,
i
Q

= force effect,

= resistance
factor,
i

= load factor corresponding to
i
Q
,

= a factor relati
ng to ductility, redundancy, and
operational importance.

The main difference between current LRFD method and nonlinear inelastic analysis method is that the
right side of Eq. (
1
)
,

(
n
R

) in the LRFD method is the resistance or stre
ngth of the component of a
structural system, but in the nonlinear inelastic analysis method, it represents the resistance or the

17

-
carrying capacity of the whole structural system. In the nonlinear inelastic analysis method, the
-
carrying capacity

is obtained from applying incremental loads until a structural system reaches
its strength limit state such as yielding or buckling. The left
-
hand side of Eq.
(1
), (
i i
Q
 

)
represents the member forces in the LRFD method, but the app
lied load on the structural system in the
nonlinear inelastic analysis method.

4
.
2

Modeling
C
onsideration

4
.
2
.1 Sections

The A
ISC
-
LRFD Specification uses only one column curve for rolled and welded sections of
W, WT, and HP shapes, pipe, and structur
al tubing

(AISC, 1994)
. The Specification also uses same
interaction equations for doubly and singly symmetric members including W, WT, and HP shapes,
pipe and structural tubing
, even though the interaction equations were developed on the basis of W
shape
s by Kanchanalai (1977).

The proposed analysis was developed by calibration with the LRFD column curve. To this
end, it is concluded that the proposed methods can be used for various rolled and welded sections
including W, WT, and HP shapes, pipe, and s
tructural tubing without further modifications.

4.2
.2 Structural members

An important consideration in making this nonlinear inelastic analysis practical is the
required number of elements for a member in order to predict realistically the behavior of
frames. A
sensitivity study of nonlinear inelastic analysis
for two
-
dimensional frames was
performed on the
required number of element
s

(Kim and Chen, 1998)
. Two
-
element model adequately predict
ed

the

18

strength of a
two
-
dimensional
member.

This rule may be
used for modeling a three
-
dimensional
member.

4.2.3 Geometric imperfection

The magnitudes of geometric imperfections are selected as
2 1,000

for unbraced
frames and
1 1,000

for braced frames.
To model a parabolic

out
-
of
-
straightness in the member
,
two
-
element model with maximum initial deflection at the mid
-
height of a member adequately
captures imperfection effects. It

is
concluded that practical nonlinear inelastic analysis is
computationally efficient. The pat
tern of geometric
imperfection
s is assumed to be the same as the
elastic first order deflected shape.

4.2
.
4

In the proposed nonlinear inelastic analysis, the gravity and lateral loads should be applied
simultaneously, since

it does not account for unloading. As a result, the method under
-
predicts the
strength of frames subjected to sequential loads, large gravity loads first and then lateral loads. It is,
however, justified for the practical design since the development of

the LRFD interaction equations
was also based on strength curves subjected to simultaneous loading and the current LRFD elastic

It is necessary, in an non
linear inelastic analysis, to input each increment load (not the total
-
displacement behavior. The incremental loading process can be
achieved by scaling down the combined factored loads by a number between 20 and 50. For a

19

h
ighly redundant structure, dividing by about 20 is recommended and for a nearly statically
determinate structure, the incremental load may be factored down by 50. One may choose a number
between 20 and 50 to reflect the redundancy of a particular structure
.

Since a highly redundant
structure has the potential to form many plastic hinges and the applied load (i.e. the smaller scaling
number) may be used.

4.3

Design Consideration

4
-
carrying capacity

T
he elastic analysis method does not capture t
he inelastic redistribution of internal forces
throughout a structural system,
since
the

first
-
order forces, even with the
1
B

and
2
B

factors,
account for the second
-
order geometric effect but not the inelas
tic redistributions of internal forces.
T
he method may provide a conservative estimation of the ultimate load
-
carrying capacity. Nonlinear
inelastic analysis, however, directly considers force redistribution due to material yielding and thus
allows smaller

member sizes to be selected. This is particularly beneficial in highly indeterminate
steel frames. Because consideration at force redistribution may not always be desirable, the two
approaches (including and excluding inelastic force redistribution) can

be used. First, the load
-
carrying capacity, including the effect of inelastic force redistribution, is obtained from the final
loading step (limit state) given by the computer program. Secondly, the load
-
carrying capacity
without the inelastic force red
istribution is obtained by extracting that force sustained when the first
member yield or buckled. Generally, nonlinear inelastic analysis predicts the same member size as the
LRFD method when force redistribution is not considered.

4.3.2 Resistance fac
tor

20

AISC
-
LRFD specifies the resistance factors of 0.85 and 0.9 for axial and flexural strength of
a member, respectively. The proposed method uses a system
-
level resistance which is different from
AISC
-
LRFD specification using member level resistance fa
ctors. When a structural system collapses
by forming plastic mechanism, the resistance factor of 0.9 is used since the dominent behavior is
flexure. When a structural system collapses by member buckling, the resistance factor of 0.85 is used
since the domi
nent behavior is compression.

4
.3.
3

Serviceability limit

According to the ASCE Ad Hoc Committee on Serviceability report (Ad Hoc Committee,
1986), the normally accepted range of overall drift limits for building is
1 750

to
1 250

times the
building height,
H
, with a typical value of
400
H
. The general limits on the interstory drift are
1 500

to
1 200

times the story heig
ht. Based on the studies by the Ad Hoc Committee (1986),
and

by Ellingwood (1989), the deflection limits for girder and story are selected as

Floor girder live load deflection :
360
H

Roof girder deflection :
240
H

Lateral drift :
400
H

Interstory drift :
300
H

At service load levels, no plastic hinges are allowed to occur in order to avoid permanent
deformations under service loads
.

4.3.4 Ductility requirement

Adequate rotation capacity is required for members to develop their full plastic moment
capacity. This is achieved when members are adequately braced and their cross
-
sections are compact.

21

The limits for lateral unbraced le
ngths and compact sections are explicitly defined in AISC
-
LRFD
(1994).

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Abdel
-
Ghaffar, M., White, D. W., and Chen, W. F. (1991). “Simplified second
-
order inelastic
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.
” Special Volume of Session on Approxi
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Verification Procedures of Structural Analysis and Design
,

Proceedings at Structures Congress 91,
ASCE, New York, 47
-
62.

Ad Hoc Committee on Serviceability, Structural serviceability (1986). A critical appraisal and
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2664.

Al
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Mashary, F. and Chen, W. F. (1991). “Simplified second
-
order inelastic analysis for steel frames.”
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399.

AISC (1994). Load and Resistance Factor Design Specification, American Institu
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Alvarez, R. J. and Birnstiel, C. (1967). “Elasto
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plastic analysis of plane rigid frames, school of
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22

Attala, M. N., Deierlei
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plastic
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hinge
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2473.

Avery, P. and Mahendran, M. (2000). “Large
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scale testing of steel frame structures comprising non
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936.

Chen, W. F. and Atsuta, T. (1977). “Theory of beam
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columns, vol. 2, space behavior and design.”
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-
Hill, New York, 732 pp.

Chen W.F. and Kim, S. E.(1997).

LRFD steel design using advanced analysis.

, CRC Press, Boca
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a.

Chen, W.F. and Lui, E. M.(1986).

Structural stability
-
theory and implementation.

Elsevier, New
York, 490pp.

Clarke, M. J., Bridge, R. Q., Hancock, G. J., and Trahair, N. S. (1992). benchmarking and verification
of second
-
order elastic and inelastic f
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CSA (19
89
). Limit States Design of Steel
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M
89
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S16.1
-
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Ellingwood (1989).

Limit states design of steel structures.

, AISC Engineering Journal,
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-
8.

EC3 (1990). Design of Steel Structures: Part I
-

General Rules and Rules for Buildings, Vol. 1,
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ECCS (1984). Ultimate Limit State Calculations of Sway Frames with Rigid Joints, Technical
Committee 8
-

S
tructural Stability Technical Working Group 8.2
-

System, Publication No. 33, 20 pp.

ECCS (1991). Essentials of Eurocode 3 Design Manual for Steel Structures in Buildings, ECCS
-
Advisory Committee 5, No. 65, 60 pp.

23

El
-
Zanaty, M., Murray, D., and Bjorhovde,
R. (1980). “
I
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Kanchanalai, T. (1977). “
T
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-
columns in unbraced steel frames.” AISI
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Texas at
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-
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1069.

Kim, S. E. (1996). “Practical advanced analysis for steel frame design.” Ph
.
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Kim, S. E. and Chen, W. F. (1996). “Practical advanced analysis for steel frame design.” The ASCE
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IL,April, 19
-
30.

Kim, S.E. and Chen, W.F. (1996a) "Practical advanced analysis for braced steel frame design", ASCE
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-
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Kim, S.E. and Chen, W.F. (1996b) "Practical advanced analysis for unbraced steel frame
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ASCE J. Struct. Eng., ASCE, 122(11): 1259
-
1265.

Kim, S.E. and Chen, W.F. (199
7
) "
Further studies of practical advanced analysis for weak
-
axis
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",
Engrg. Struct., Elsevier
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9
(
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407
-
416
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Kim, S.E. and Chen, W.F. (1998)
.

"A sensitivity study

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-
hinge
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-
673.

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-
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2
5.

Kim, S.E. and Choi, S.H.(2001). "Practical advanced analysis for semi
-
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,

Solids
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51), 9111
-
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24

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, Kim, Y. and

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Nonlinear analysis of 3
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walled
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Large
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-
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-
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-
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-
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Liew, J.Y.R
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25

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SS
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Plastic hinge methods for advanced analysis of steel frames.

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Ziemian, R. D.(1990).

Advanced methods of inelastic analysis in the limit states design of st
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, Ph.D. Dissertation, School of Civil and Environmental Engineering, Cornell University,
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26

Ziemian, R. D., White, D.W., Deierlein, G. G., and Mcquire, W.(1990).

One approach to inelastic
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, AISC, Chicago, 19.1
-
19.

TABLE 1.

Summary of Tension Coupon Tests

Section

Member

number

Specimen

y

ksi

*
y

10
-
5

st

10
-
5

E
st

ksi

ult

ksi

Elongation

in 8 in,

W8

17

(A36
-
70A)

C1A

C1C

Flange

37.9

128

1140

442

62.4

28.2

Flange

37.7

127

1378

356

-

29.7

Web

40.6

137

2450

345

61.7

32.9

M4

13

(A572
-

73)

B1,B2

B3,B4

Flange

48.5

164

1203

406

69.6

26.6

Flange

48.6

164

1062

399

69.9

27.2

Web

50.1

169

2228

323

69.5

26.7

C1B and C2B were no
t tested

*
y
=

y
/E(E=29,500ksi)

TABLE 2. Measured Properties of Beam and Column Section

Frame

Section

Handbook

EI

(kip
-
in
2

10
4
)

Measured

EI

(kip
-
in
2

10
4
)

Handbook

M
p

(kip
-
in)

Measured

M
P

(kip
-
in)

C

12B16.5

310

271

742

845

C

10B15

203

190

576

635

C

6WF15

158

165

686

760

TABLE 3. Dimensions and Properties of Members

Test
Section

D

b
r

t
r

t
w

r
1

A
g

I

S

y

27

frame

(mm)

(mm)

(mm)

(mm)

(mm)

(mm
2
)

(10
6
mm
4
)

(10
4
mm
3
)

Flange Web

2

310UB32

298

149

8.0

5.5

13.0

4080

63.2

475

360 395

TABLE 4. Actual Se
ction Properties of One
-
Quarter

Scaled Frames

A

(cm
2
)

I

(cm
4
)

Z

(cm
3
)

Z
p

(cm
3
)

y

(t/cm
2
)

Column

21.8

391

77.4

88.5

2.64

Beam

10.6

177

35.0

40.6

3.04

TABLE 5. Dimensions and Material Properties of Equilateral Triangular Space Frame

L

(in)

D

(in)

T

(in)

E

(ksi)

G

(ksi)

y
(ksi)

Column Beam

All members

48

1.682

0.176

28800

11520

30.6 31.1

TABLE 6. Dimensions and
Properties

of Section H
-
150

150

7

10 Used in the Frame

H
eight

mm
H

Width

mm
B

Thickness

of Flange

mm
t
f

Thickness

of Web

mm
t
w

Fillet

mm
r
1

Axial
Area

2
mm
A
g

Moment of
Axis

4
6
10
mm
I
X

Moment of
Axis

4
6
10
mm
I
Y

Nominal

150

150

10

7

11

4014

16.40

5.63

Measured

Column

152.3

149.9

10.2

6.75

-

4053

17.20

5.74

Beam

149.1

150.0

9.2

6.50

-

3713

15.14

5.18

TABLE 7. Comparison of Experime
ntal and Design Load Carrying Capacity

(a) Experiment

(b) Analysis

(c) AISC
-
LRFD design

(b)/(a)

(c)
/(a)

P

612.0

612.0

443.5

1.0000

0.7247

H

169.2

175.5

122.6

1.0372

0.7246

TABLE
8
. Comparison of Experimental and Design Load Carrying Capacities

28

(a) Experiment

(b) Analysis

(c) AISC
-
LRFD design

(b)/(a)

(c)/(a)

Test frame
3

P

681.8

680.9

510.2

0.9985

0.7483

H
1

136.4

136.2

1 0 2.0

0.9 9 8 4

0.7 4 8 1

H
2

6 7.5

6 8.1

5 1.0

1.0 0 8 3

0.7 5 5 6

F I G. 1. A
n a l y s i s a n d De s i g n Me t h o d

29

FIG. 2. Interaction
between

A Structural System and Its Component Members

-
Deformation Characteristics of Plastic Analysis Methods

30

FIG. 4. Concept of Spread of Plasticity for Various Advanced A
nalysis Methods

FIG. 5. Model of Plastic
-
Zone Analysis

31

FIG. 6. Explicit Imperfection Model for Elastic
-
Plastic Analysis Recommended By ECCS

FIG. 7. Examples on Application of Notional Loads for Second
-
Order Elastic
-
Plasic Hinge
Analysis

32

F
IG. 8. Two
-
Bay Frame

FIG. 9. Axial Load
-
Deflection Behavior of Specimen

33

FIG. 10. Specimen for Three
-
Story Frame

FIG. 11. Lateral Load
-
Sway Behaviour of Frame C

34

FIG. 12. Schematic Diagram of Test Arrangement

FIG. 13. Sway Load
-
Deflection Curve for Test Frame 2

35

FIG. 14. One
-
Quarter Scaled Frames.(From Wakabayashi, M. And

Matsui, C., Trans. Arch. Inst. Jpn. 193,17,1972, With Permission
)

36

FIG. 15. Horizontal Force
-
Displacement Behaviours o
f One
-
Quarter Scaled
Frame.(Two Story).(From Wakabayashi, M. And Matsui, C., Trans.

Arch.Inst. Jpn. 193,17,1972, With Permission)

37

FIG. 16. Harrison

s Space Frame(Harrison 1964)

-
Deflection for Harrison

s Space Fra
me

38

H2

H1

Base
2nd floor
P
P
P
Roof
P
2.20m
1.76m
2.5m
3.0m
X
Z
Y

FIG. 18. Dimension and Loading Condition of Test Frame

X
Z
2.20m
3.0m

2.5m

Y

Base
2nd floor
1.76m
P
P
P
Roof
P
H

FIG. 19. Dimensions and Loading Conditions of Test Frame in Main Test

39

0
10
20
30
40
H
o
r
i
z
o
n
t
a
l

d
i
s
p
l
a
c
e
m
e
n
t

(
m
m
)
0
40
80
120
160
200
H
o
r
i
z
o
n
t
a
l

l
o
a
d

(
k
N
)
Experiment(H1)
Analysis(H1)
Experiment(H2)
Analysis(H2)

F
IG
.
20
.

Comparison of
Horizontal

L
-
Displacement

C
urves for
Space T
est
F
rame

2

F
IG
.
21
.