EFFECTS OF EARTHQUAKE EXCITATIONS ON RETICULATED DOMES/ · 2020-01-20 · THE\EFFECTS OF EARTHQUAKE...
Transcript of EFFECTS OF EARTHQUAKE EXCITATIONS ON RETICULATED DOMES/ · 2020-01-20 · THE\EFFECTS OF EARTHQUAKE...
THE\EFFECTS OF EARTHQUAKE EXCITATIONS ON RETICULATED DOMES/
by
David A.“Uliana„
Thesis submitted to the Faculty of theU l
Virginia Polytechnic Institute and State University
in partial fulfillment of the requirements for the degree of
MASTER OF SCIENCE
inU CIVIL ENGINEERING
APPROVED :
D . S. M. HolzerD
Dr. A; E. Somers, Jr. Dr. R. H. Plaut
February, 1985Blacksburg, Virginia .
1/6E
ACKNOWLEDGMENTS y:2éär The author would like to express his gratitude to Dr. S.\‘ i
M. Holzer for his guidance, unlimited patients, and helpful
sugestions throughout the researching and writing of this
manuscript. ·The completion of this thesis would have been
impossible without his help. Gratitude is also extended to .
Dr. A. E. Somers, Jr., and Dr. R. H. Plaut for their help-
ful comments and for reviewing this thesis.
A special and most sincere thanks is extended to David,
Gloria and Morgan Young for being my family away from home,
to Mun Fong Chan for being an exceptional officemate and
consultant, and to Annie Crate and John Volk for their
friendship through the difficult times.
The financial support provided by the Department of Civil\\‘\~
Engineering and the use of the Virginia Tech computing fa-‘
‘—~a.. cilities are also greatly appreciated.
ii
‘ TABLE OF CONTENTS
Q Q Q Q Q Q Q Q Q Q Q Q Q Q Q Q Q Q i i
Q Q Q Q Q Q Q Q Q Q Q Q Q Q Q Q Q Q V
-
Q Q Q Q Q Q Q Q Q Q Q Q Q Q Q Q Q
QChapter· Page
I Q Q Q Q Q Q Q Q Q Q Q Q Q Q Q Q Q Q 1
Objectives ................. 1Q Q Q Q Q Q Q Q Q Q Q Q Q 3
II. STRUCTURAL MODEL ............... 8
Introduction .............3. . . E 8The Tangent Stiffness Matrix ........ 9Dome Models ................ 12
III; Q Q Q Q Q Q Q Q Q Q Q Q Q Q
QIntroduction................ 17Newton-Raphson Method ........... 20
_ Convergence Criteria ............ 23Discussion of Stability .......... 26
IV Q Q Q Q Q Q Q Q Q Q Q Q Q Q Q
QIntroduction................ 31Modeling Earthquake Excitations ...... 32Damping Matrix ............... 34The Newmark-Beta Method .......... 36
Linear Analysis ............. 36Nonlinear Analysis ............ 37
Discussion of Time Steps .......... 39
V Q Q Q Q Q Q Q Q Q Q Q Q Q Q Q Q Q Q Q Q 4 1
Q The Finite Element Program ......... 42Comparison with Previous Results ..... 42Checks on Some Assumptions ........ 58 Q
Small Lamella Dome ............. 59Static Analyses ............. 59Dynamic Analyses ............. 66
Large Dome ................. 68Static Analyses ............. 68
4 Dynamic Analyses ............. 83 _
iii
‘ TABLE OF CONTENTS (continued)' ESS
Comparisons of Stresses for Both Large andSmall Domes ............... 90 U
VI. CONCLUSIONS AND RECOMMENDATIONS ....... 100
Conclusions ................ 100‘ Recommendations for Future Research .... 101
REFERENCES .................... 104
Aggendix gage” A. FREQUENCIES OF VIBRATION FOR THE SMALL LAMELLA
DOME ................... 110
B. FREQUENCIES OF VIBRATION FOR THE LARGE LAMELLADOME ................... 111
C. PROGRAM LISTING ............... 114
VITA . . .......... . .......... 189 ·
ABSTRACT
iv
LIST OF TABLES
Table pass1. Maximum stresses of static analysis for numbered
members of the small dome. ........... 63
2. Maximum allowable stresses and buckling stresses fornumbered members of the small dome........ 64
3. A comparison of maximum stresses in numbered members _of the small dome for dead load only....... 75
4. A comparison of maximum stresses in numbered membersof the small dome for dead load plus 10 lb./ft.2. 76
5. A comparison of maximum stresses in numbered members· of the small dome for dead load plus 20 lb./ft.2. 77
6. A comparison of maximum stresses in numbered membersof the small dome for dead load plus 30 lb./ft.2. 78
7. A comparison of maximum stresses in numbered membersof the small dome for dead load plus 40 lb./ft.2. 79
8. Maximum stresses of static analysis for numberedmembers of the large dome. ........... 82
9. Maximum allowable stresses and buckling stresses fornumbered members of the large dome........ 84
10. A comparison of maximum stresses in numbered membersof the large dome for dead load only....... 92
11. A comparison of maximum stresses in numbered membeisof the large dome for dead load plus 20 lb./ft. . 94
12. A comparison of maximum stresses in numbered membegs yof the large dome for dead load plus 40 lb./ft. . 96
v
LIST OF FIGURES
Eigum pass1. Member end displacements for a typical truss
element. .................... 10
2. Large lamella dome. ............... . 13
3. Small lamella dome ................. 15
4. The tributary area for the top node of the domes. . 19
5. The Newton—Raphson method for a one degree—of-freedom system.................. 22
6. The Modified Newton-Raphson method for a one degree-_ of-freedom system. ............... 24
7. Equilibrium path for a one degree-of-freedom system. 29
8. One degree—of-freedom plane truss presented byNickell. .................... 43
9. Displacement vs. load graph for the static analysislof the one degree-of-freedom truss presented by
. Nickell. .................... 44
10. Displacement vs. time graph for the dynamic analysisof the one degree—of-freedom truss presented byNickell. .................... 45
·_ 11. Ten member plane truss presented by Noor and Peters 46
12. Displacement vs. time graphs of joint A for lineary and nonlinear analyses of the ten member truss. . 48
13. Axial force in member B for the linear analysis ofthe ten member truss............... 49
14. Axial force in member B for nonlinear analysis ofthe ten member truss............... 50 _
15. A comparison of the stress-strain relationships. . . 51
16. Large ten-bay truss presented by Noor and Peters. . 53
viM
17. The displacement of joint A for the linear analysisof the ten-bay space truss............ 54
18. Axial force in member B for linear analysis of theten-bay space truss. .............. 55
19. Displacement of joint A for nonlinear analysis ofthe ten-bay space truss. ............ 56
20. Axial force in member B for nonlinear analysis ofthe ten-bay space truss. ............ 57
21. Small lamella dome with numbered members. ..... 60
22. Load vs. displacement curves for uniform load on thesmall dome....4................ 61 B
23. Load vs. displacement curves for nonuniform load onthe small dome.................. 65
24. Accelerations, velocities and displacements for theEl Centro earthquake of May 18, 1940....... 67
25. Displacement of joint A for dead load only on thesmall dome.................... 69
26. Displacement of joint A for dead load plus 10lb./ft. on the small dome............ 70
27. Displacement of joint A for dead load plus 20 lb./ft.2 on the small dome...........·. . . 71
28. Displacemegt of joint A for dead load plus 30lb./ft. on the small dome............ 72
29. Displacement of joint A for dead load plus 40lb./ft.2on the small dome. ........... 73
30. A comparison of the displacements of node A forseveral live loads on the small dome....... 74
31. Large lamella dome with numbered members. ..... 80
32. Load vs. displacement curves for uniform load on the .large dome.................... 81
33. Load-displacement curves for nonuniform load on thelarge dome.................... 86
viil
34. Displacement of joint A for dead load only on thelarge dome.................... 87
35. Displacemegt of joint A for dead load plus 20lb./ft. on the large dome............ 88
36. Displacement of joint A for dead load plus 40lb./ft.2 on the large dome............ 89
37. A comparison of the displacements of joint A for thetwo live loads on the large dome......... 91
xdii
Chapter I
INTRODUCTION
1-1Dome structures, in particular reticulated domes, have
demonstrated their usefulness as strong, light-weight cover-
ings over large areas such as stadiums, industrial buildingsÜ
and exhibition centers where large, unobstructed areas are-
required. They enclose a maximum amount of space with a
minimum surface and are economical because they require lit-
tle material as compared to other structures. Several
different dome configurations exist. Some of the more popu-
lar configurations are the lamella, the geodesic, and the
Schwedler domes.· —
Many studies have been performed by various researchers
to analyze these domes under static loads. They have re-
peatedly demonstrated the domes' strength and economy for
Civil Engineering applications. The author found no refer-
ences, however, that presented results for dynamic analyses
of full sized reticulated domes. Due to an increasing in-
terest in earthquake resistant design of structures, dynam-
ic analyses of these domes are definitely needed. .
This thesis is a study of the behavior of full sizedt
domes, designed under static criteria and subjected to
2
earthquake excitations. Static analyses are performed on
two lamella domes. The axial stresses and strains in all
members are determined for several distributed static loads.
Dynamic analyses are then performed on the same structures
with the same distributed loads. The dynamic loads are in
the form of earthquake excitations to the bases of the
domes. Axial stresses and strains in all members are deter-
mined and compared to those of the static analyses. The
displacements of several nodes for both analyses are given
and effects of the earthquake motions on the stability of
the domes and their members are determined.
A finite element computer program written by the author
(see Appendix C) was used to perform these analyses. The
program can analyze geometrically and materially nonlinear
behavior of space trusses subjected to static or dynamic
loads. The direct integration procedure used is the New-
mark-Beta method and the nonlinear solution technique is the
Newton-Raphson method. The program can calculate the dis-
placements, velocities, and accelerations of all nodes as
well as the axial stresses and strains of all members at
each load or time step. The ability to determine if any
member reaches its buckling load or if the dome becomes glo-
bally unstable is also present in the program._
1 l
3l-2LIIEBAIHBE BEYLEHU
2A literature search revealed that there is little re-
search done on the dynamic analysis of space truss struc-
tures. No references were found that addressed the topic of2 earthquake excitations on reticulated domes. Severalrefer-ences
were found that closely relate to this topic, though
none use the space truss specifically. Penzien and Liu (43)
and Tanabashi (52) address the solution of nonlinear struc-
tures subjected to earthquake excitations while Berger and
~ Shore (8) and Carter, Dib and Mindle (10) present the dynam-
ic response of nonlinear structures.
In order to cover the thesis topic more carefully, three
subjects were searched for pertinent literature. These sub-
jects are:
1. Analysis of space trusses
2. Methods for structural dynamics solutionsl
3. Modeling earthquake excitations·
A review of these subjects provides a review of the thesis
topic.
One method of analysis that is used to model large space
trusses, such as reticulated domes, is to replace the truss
with a continuum model. The continuum model may then be j
solved with familiar methods of mechanics. Forman andI
Hutchinson (16) and Crooker and Buchert (12) present methods
4
to replace a space truss with a continuous shell. Renton
(45) models a repetitive space truss as a continuous beam
and Noor, Anderson and Greene (40) offer approximations for
beam-like and plate-like lattices. Conventional finite ele-
ment analyses of large systems would be expensive and time
iconsuming. These methods allow relatively inexpensive anal-
yses of large repetitive structures.”
Force and displacement methods are popular for analyzing
(trusses since a truss model may be defined by a finite num-
ber of forces or displacements. Methods such as Maxwell-
Mohr and matrix analysis are relatively simple for linear
models. Goldberg and Richard (19) and Hensley and Azar (22)
derive analyses methods for material nonlinearities while
Epstein and Tene (15) and Jagannathan and Epstein (29) con-
sider the geometric nonlinearities. A mixed formulation is
provided by Noor (38) who considered both material and geo-
metric nonlinearities, and a geometrically nonlinear stiff-
ness matrix for a space truss element is derived by Baron
and Venkatesan (2) with perturbation techniques. Reducing
the number of degrees-of-freedom for these nonlinear prob-° lems is desirable for many analyses. A discussion on the
reduction of the number of degrees-of-freedom of nonlinear
problems is given by Noor (39).l
5 g
For trusses that should be modeled with imperfect mem-
bers, Rosen and Schmit (47) and (48) have both an accurate
and an approximate method to include the imperfections. If
the effects of transverse loads are desired, then a frame
element is needed. A model for a geometrically nonlinear
pin-jointed frame is presented by McConnel and Klimke (33).
The pin-jointed frame has the characteristics of a truss but
permits bending strains. For dynamic analyses of nonlinear .
_ space trusses, which is required for this study, Noor and
Peters (41) use a mixed method.
Several techniques exist for the solution of structural
„dynamics problems. Among some of the less familiar methods
are the Finite Strip Method (11), the Runge-Kutta Method
(27) .and the Fast Fourier Transform Method (20). These
methods have not proved to be as popular for solutions of
multi—degree-of-freedom systems as modal analysis and direct
integration. Der Kiureghian (13) presents a probabilistic
approach to linear modal analysis and Nickell (36) presents
a deterministic approach to nonlinear modal analysis.
For solution of nonlinear problems like the space truss,
direct integration proves to be much more suitable than mo-
dal analysis, according to Ref. 18. Several different di-
rect integration procedures exist. Some of these methods
are the Central Difference method (42), the Wilson-6 method
6
(54), the Newmark method (34) and the method proposed by
Ginsburg and Gilbert (18). Two important factors to consid-
er when using direct integration are the stability and ac-
, curacy of the solution. Nickell (35), Preumont (44) and
Bathe and Wilson (7) discuss stability and accuracy while
Hinton, Rock and Zienkiewicz (24) and Bathe and Cimento (6)
present methods for increasing the accuracy and economy of a
solution. Atalik (1) makes solutions more economical by
constructing linear approximations for the element mass,
damping and stiffness matricies for nonlinear problems.
Earthquake excitations may be modeled in two ways. Eith-
er an actual digitized accelerogram may be used or an arti-
ficial accelerogram may be produced. Real digitized data,
such as that given by Haroun (21), are available for many
earthquakes. Data in this form, however, may not be useful
if it is digitized in time increments that are not appropri-
_ ate. Liou (32) offers a numerical method for redigitizing
seismic data without producing superficial frequencies.
An accelerogram may be generated artificially if it is
more desirable than the available digitized data. Earlier
methods for producing artificial accelerograms were mostly
deterministic. Bernrueter (9) finds maximum acceleration
only, St. Balan, et. al. (50) generate an accelerogram as aW
series of trigonometric functions and Wong and Trifunac (55)
7
0 use a Fourier transform method. These and other determinis-
tic methods have been replaced by probabilistic formulations
due to an increasing interest and understanding of the vari-
ability of earthquake ground motions._ Some of the most re—
cent research on this topic is presented by Kubo and Penzien
(31), Kameda and Ang (30), Housner and Jennings (27), Trifu-
nac (53) and Iyengar and Iyengar (28). These studies2
demonstrate how to generate artificial accelerograms through
probabilistic methods.
Because of the distinct advantages of reticulated domes
and the increasing interest in earthquake resistant design
of structures, the thesis topic should be of increasing in-
. terest in the future.•
Chapter II
STRUCTURAL MODELy
In this chapter the tangent stiffness matrix for a geome-
trically nonlinear space truss element is presented. It is
derived by the familiar principle of virtual work. This
y chapter also presents the domed structures that will be used
in the analysis. p
2.1Areticulated dome is a collection of one-dimensional
elements assembled in space so that their joints lie on a
spherical surface. These one-dimensional elements may be
modeled either as truss elements, where only axial elonga-
tion and contraction are considered, or frame elements, for
which torsion and bending about two axes are also consid-
ered. For a space truss each element has six translational. degrees of freedom. For a space frame each element has six
translational and six rotational degrees of freedom.
A space truss element was chosen for this study for the
following reasons:
l. Members of a reticulated dome are usually joined in a
fashion that more closely resembles a pinned connec-N
tion than a rigid connection.
“ 8
9
2. Dynamic nonlinear analyses are, by nature, time con-
suming. The space truss element has fewer degrees of
freedom and is more economical.
The geometry of a reticulated dome and slenderness of its
members allow for the possibility of large displacements and
rotations although the strains that develop can be small.
This results in geometrically nonlinear behavior which is
the result of a nonlinear strain-displacement relation and
equilibrium being formulated on the deformed structure. For
these problems strains are not generally proportional to
displacements. The nonlinearity must be considered in the
formulation of the stiffness matrix. .
92.2 IHEIAblQELNIMAIE.lX p
The tangent ustiffness matrix used in this study was
derived by Holzer and is presented in detail in Ref. 25.
The undeformed state and the deformed state of a typical
truss element are shown in Fig. l. During the deformation
the element may translate and rotate significantly with only
small strains being produced. Expressing the displacements
of the element as the sum of rigid body displacements and
deformations is desirable. The initial, undeformed state is
labeled I.S., the rigid body state is labeled R.B.S. and thei
deformed state D.S. The tangent stiffness matrix of an ele-
ment i is given by F
1i 10
2
Z: == == ——
. / 6, / 'd dz /
_ //.5. d
d 54
l
Figure 1: Member end displacements for a typical truss ' l
element.
I
11
. 2l _ 8 Ü 1. K · tzzrzzü ( ’. J k
where U is the strain energy of the element. The resulting
matrix is the sum of the familiar linear stiffness matrixfor a space truss and a nonlinear matrix. They combine to
form the tangent stiffness matrix for an element given by
. zi --1.K1 L _ K_ (2). _gl il
where
L. 7 c c c cLO_ 1 1 2 1 3· Fi = Y-
1 —Ül (1 CK) C (3)C 1 L 2 2C3O1 L. 7svm -——i -(l-C“)
-· ‘ _ L 3O.1.
and I O
- EiAi (4)Y Yi = T7l
For each member 'i', Ej_is Young's Modulus of the materi-al, A_ is the cross-sectional area, Li is the deformedlength and L0_is the undeformed length. Cl, C2, and C3 are
lthe direction cosines in the global 1, 2, and 3 directions,
respectively.O _
~ 12
2-3 DQME MQDEL§(
Two structural models are used for analysis. One dome is
a large lamella dome used in a study by Richter (46). A di-
agram appears in Fig. 2. The dome has a 100 foot base, a 14
foot height at the center and lies on the surface of a 200
foot diameter sphere. All of the members are tubular with
an outside diameter of 6.00 in., a cross sectional area of
3.18 in.2, and a cross sectional inertia of 15.64 in.4. The
material is T60—6011 aluminum (49) with a yield strength of
·35 ksi. and a modulus of elasticity of 10,300 ksi. The la-
mella dome has three rings and ten radial sections. This
structure was chosen for several reasons:
1. Data for the static analysis is readily available.
2. The geometry of the lamella dome is relatively easy
to define.
3. There is a natural truncation ring at the base of the
dome.Two
different boundary conditions are used for this dome,
one for the static analysis and one for the dynamic analy-
sis. For the static analysis a complete dome with a tension
ring provides the most realistic model. With the tension
ring the dome has 150 members, 61 joints, and 149 degrees of
freedom. It is constrained in the vertical direction at ev-T
·ery base joint, in both horizontal directions at the top of
I
14
the dome (to avoid translation of the dome) and in two other‘ _ horizontal directions (to avoid rotation of the dome). For
the dynamic analysis base joints are pinned. This model has
120 members, 61 joints, and 93 degrees of freedom. Since
each base joint is constrained in all three directions the
tension ring is rendered useless and the members are not in-i cluded.
For the second model a smaller lamella dome was desirable
because:
1. A comparison between similar domes of different sizes ·O
could be made.
2. Fewer degrees of freedom will require far less compu-
ter time; therefore, more extensive testing may be
performed.
The dome chosen was a truncation of the larger dome at its
second ring (See Fig. 3). All element properties and dimen-
sions are the same as in the larger dome.
As before, two different boundary conditions for this
dome are considered. For the static analysis the dome with
n the tension ring has 70 members, 31 joints, and 69 degrees
of freedom. It is constrained in the same manner as the1
«
larger dome. The fixed base dynamic model has 50 members,
1
p 16 T
31 joints, and 33 degrees of freedom. This 33 degree-of-
freedom dome requires much less computer time per run and is
used in a majority of the dynamic testing.
A geodesic geometry was considered for comparison with
the small lamella dome. However, a small geodesic with the
y same number of joints and the same overall dimensions as the
lamella could not be defined without severely modifying the
geodesic geometry to form a base ring. The geodesic geome-
try has no natural truncation ring for a shallow dome.
Also, a limit on the expense of the dynamic analyses was de-O
sired, therefore, the geodesic dome was not included.
Chapter III
STATIC ANALYSIS
In this chapter a description of the static analysis pro-
cess is given. The nonlinear solution technique used is theNewton-Raphson method. This description is followed by dis-
cussions on convergence criteria and stability of domed
structures.
3-1Adescription of the static analysis process is included!
in this study. The reasons that it is included are:‘ 1. Static analyses are the most common analyses for
space structures.
2. The dynamic analysis is actually reduced to an equi-1
valent static analysis at every time step by the New-
mark-Beta method.
3. A static analysis should always precede a dynamic· analysis when designing a structure.
' The static procedure was used in this study to determine
the nodal displacements and the element stresses caused by
several distributed live loads. The live loads used are 10, 7
20, 30, and 40 lb./ft}!. These values represent the maximumV
snow loads for most regions of the United States according
17
1 18
to Ref. 17. Several loads are used in order to show a com-
parison between loads, displacements and stresses.
The distributed load must be transformed into a series of Yconcentrated nodal forces in order to include it in the ana-
lysis. Each force is equal to the snow load times the
tributary area around the node. An example of the tributary
area for the top node of both domes is given in Fig. 4.
This area is the projected area on the horizontal plane and
is equal to 250.8 square feet. The tributary area for each
node in the first ring (node type A of Fig. 2) is 154.2
square feet. The tributary area for each node of the second
ring is 145.8 square feet for node type B of Fig. 2, and
149.1 square feet for node type C. The same areas apply to. the small dome of Fig. 3. A load vector may be formulated
by multiplying the distributed load times these areas.
Transforming the distributed load into a load vector is a
realistic assumption for many domes since suspending a load
carrying sheet from the joints of a dome is common. For
these domes the snow load is truly applied to the nodes.
For many other domes, however, rigid panels are attached di-
rectly to the members. In these domes bending strains de-
velop that cannot be neglected. A paper by Renton (45) ad-
. dresses the modeling of space trusses for beam-like behavior4
and McConnel and Klimke (33) discuss the analysis of pinned-
919
1
/°’ ·“
._•
,K VV [
„: ··‘“ A ¢•‘Z«;%§%•:••*11{~
°RRo„1EoT1oHow THE HoR1ZoHTALREAHE FORTOP@—TRlBUTARY AREA. (g5o.6 SEH.)
Figure 4: The tributary area for the top node of the domes.
20
end space frames. In this study, however, the bending ef-
' fects are not considered.
3-Z MEIHQD „The basic equation to be solved in nonlinear analysis is
. Q _ A+AAF = 0 (5)
Q is the external nodal load vector which can also be writ-
ten as (A+AAfQ where A.is a scalar andQ”is
a constant vec-
tor, and k+AMF is the nodal force vector that results from yelement stresses. When Eq. (5) is satisfied the system is
in equilibrium. A nonlinear solution technique is requiredl
since k+AÄF is not a linear function of the nodal displace-
ment vector )‘+A^q. This description of the nonlinear solu-
tion technique is given for the static analysis. The reader
should understand that the technique is the same for dynamic
analyses where t may replace A.
One of the most popular techniques for the solution of
nonlinear equations is the Newton-Raphson method. By apply-
ing this method Eq. (5) becomes
K(i)Aq(i+l) = Q _ F(i) = R(i) (6)
for the i'th iteration where K(i)is the tangent stiffnessO
matrix at q(l)- A complete derivation appears in Ref. 3.
212
Eq. (6) is solved repeatedly until a new equilibrium state
is obtained with a sufficient degree of accuracy (See sec-
tion 3.3).
The Newton-Raphson method requires a known equilibrium°(
position (qO,Q0) to start the iterative process. Refer to
Fig. 5 for a graphical representation of a one degree-of-
freedom system. The load is incremented from Q0 to Q where
a new equilibrium position is sought. The tangent stiffness
matrix K(O)at q(0)is determined and Eq. (6) is solved for
i=0 and R(0) =Q-F(0) . A new equilibriumpositionq(i+l)
2 q(i) + Aq(i+l) (7)can be found along with the nodal point forces F(i+l)
corresponding Residual forces
R(i+1) 2 (Ä 2 Mm- _ F(i+l) (8)
are determined and a convergence check is performed. If
convergence is not reached then the tangent stiffness matrixK(i) is computed and Eq. (6) is solved. This process is re-
peated until convergence is reached.
A modified method can also be used. The tangent stiff-
ness matrix K(0) is found for the first iteration only and
Eq. (6) is solved using K(0) for K(i) for every iteration.A
A graphical representation for a one degree-of—freedom sys-
zz
.l
Ko Kl.
Q :I ‘ER*· I f' II I · I
II I| I| II II ·F=I
. I: |
I |·
QO °·· ;° II „ II ; II I I
I I [Ä ld I ZX QQ ICI (LG). Figure 5: The Newton—Raphson method for a one degree-of-
~I
freedom system.
D 23
tem appears in Fig. 6. Using a constant tangent stiffness
matrix eliminates the need for the costly processes of com-
- puting and factorizing the stiffness matrix for every itera-
tion.
3-3 §.BI.IEBlAMathematically, the Newton-Raphson procedure will con-
verge to an exact solution; however, it will not do so in a
finite number of iterations. A condition referred to as aconvergence criterion must be defined so that the iterative
process may be stopped at an appropriate vector q(i+l)in Eq.
(7). The convergence criterion may be a check on the magni-
tude of Aq$i+l) or on the magnitude of the unbalanced force
vector R(i+l{ It must assure that the desired accuracy
ofthesolution is found without excessive computational ef-
fort. Once the condition is satisfied and convergence is‘ achieved the Newton-Raphson process is ended and the analy-
sis proceeds on to the next load or time step.
Several methods exist to check convergence. The·simplest
is convergence of the norm of the change in the displacement
vector °
||¤9(i+l)|| < 9
(9)wherethe norm may represent either the Euclidean norm
= Aq§j_+l)2)]-/2 (10)
’ 24
ÄK7·°Q " "'°'”'” ” ° I-. ° _-
I I. I I I
I · I. I II I I .I I I .I II I I' I I I I
I I I I
I I I I
I I I II I I I
I I III I I I ' I ‘
QI I I I I‘ "" I ' I IO I I I II
I I I I |I I I I I I‘ I I { I I
A10) I
Figure 6: The Modified Newton-Raphson method for a oneI
degree-of-freedom system.
25
‘ or the largest absolute value of the components of°Aq(i+lÄ
It is apparent that the norm is not a dimensionless quanti-
ty, therefore, the value for e must be determined by consid-
ering the dimensions of the actual system. This creates a
problem in that e may need to be changed for analyses of
various systems. To avoid this problem the nondimensional
criterion(i+l) _ (i+l)
. (1+1) .may be used. Again, ||q || is the norm of the total
displacement vector and may be either the Euclidean norm or
the largest single component of the vector. The criterion
used throughout most of this study was that of Eq. (ll).
Bathe (6) suggests using e=0.00l for this criterion.
A check on the unbalanced force vector is similar to the
methods presented above. The simpler check is on the un-
balanced force vector
('+l)l i- °
IIRL ||<eR (12)( '+l
äwhere ||R“L )|| may be either
. 1) . 2 l/2IlR(l+ ll = (Z R$l+l) ) (13)· JJ
or the largest absolute value of the vector components.l
Here, too, e may need to be changed for different systems.R
66
26
In order to nondimensionalize this quantity, the unbalanced
force vector must be divided by the restoring force vector4 to create the ratio
(14)
When this ratio is small enough then the R(i+l)vector is in-
significant and convergence is reached. This would guaran-
tee thatFLi+l)
is close enough to Q to assure convergence.
The value of eR could remain unchanged for different sys-
tems.
3-4Reticulated domes are very strong for their light weight .
and are able to resist great static loads. For example, the
large lamella dome used in this study was able to resist a4
static load of 24 times its own self-weight. The geometries
and relative lightness of these domes, though, make them
susceptible to instabilities. A reticulated dome may become
unstable long before yield stresses in the material are
reached. Special consideration of instabilities may be re-
quired in order to effectively design a reticulated dome.
For the space truss the possibilities of member instabil-
ities and global instabilities exist. For a truss member
the internal force increases linearly with axial deformation
27
in compression until it is near the buckling load. In the
vicinity of the buckling load a real, imperfect member ex-
periences transverse deflections that increase significantly
for small increases in loading provided the material remains
elastic. The truss elements used in this study, however,
are modeled as perfect elements that do not allow transverse
deflections. In order to include a more realistic behavior,I the compressive loads are assumed to increase linearly with
axial deformation until the buckling load is reached, after
which the compressive force is held constant at the buckling
load. The buckling of a member may or may not result in a
global instability that is defined below and is usually less
catastrophic. One reason for this may be that the buckling
of a member could redistribute forces in the structure andl
adjacent members may compensate for the loss in stiffness.
Global instabilities result from a decrease or loss in
stiffness of the structure and exist in two forms: limit
point instabilities and bifurcation instabilities. Either
of these can be determined mathematically using the tangent
stiffness matrix. For the nonlinear problem the tangent
stiffness matrix is updated at least once every load step
while for some analyses it is updated more often. Instabil-
ities cause the stiffness matrix to become singular. Forl
the bifurcation instability, the Newton—Raphson procedure
7 ( 28
may be used to trace beyond the point and over the unstable
post-bifurcation curve. It cannot be used, however, to
trace an equilibrium path beyond a limit point. The inabil-
ity to proceed past a limit point is an inherent weakness of
_the Newton-Raphson method. To analyze the dome past a limit
point other methods such as the Riks—Wempner method (26) are
required.
A practical description of both types of instabilities is
presented by Supple (51). He states that under increasing
deformation the dome may gradually lose its stiffness until
a complete loss of stiffness is realized. At this time a7
dynamic jump occurs to a highly displaced configuration that
·is usually symmetric. This may be a limit point instabili-
ty. A graphical representation for a one degree-of-freedom
system is shown in Fig. 7. The complete inversion of a
domed structure is an example of the equilibrium state after
a limit point instability. The dome may also exhibit a sud-
den loss of stiffness and displace into a configuration that7
is different from its initial configuration. This is the
bifurcation instability. The suddenness of the loss in
stiffness is what differentiates the bifurcation from the 7
limit point instability. The bifurcation is shown for a one
degree-of—freedom system in Fig. 7. The equilibrium con-7
figuration reached after a bifurcation instability does not
29 Vr
stable path
'K limit point -
I—~ __, __ __ __ __ new ”_= j equilibrium//’
\\F<' C) configuration.
unstable/\*
pathbifurcationpoint
:table path "
Figure 7: Equilibrium path for a one degree—of-freedomsystem.
Chapter IV
DYNAMIC ANALYSIS
In this chapter the equations associated with the dynamic
analysis are presented. The dynamic finite element equa-
tions are given first followed by the method used to include
earthquake excitations. The direct integration scheme cho-
sen for this study is the Newmark-Beta method. The method's
application to linear and nonlinear problems is given along
with the damping matrix formulation and a discussion on time
steps.
4-1„The equations of motion ufor a multi-degree-of-freedom
system can be expressed
Mä + cq + Kq = Q (15)
whereT
M-the mass matrix yC-the damping matrix
_ · K-the stiffness matrixq-the nodal acceleration vector
q-the nodal velocity vectorA
q-the nodal displacement vector
Q-the dynamic nodal load vector
31
32
The type of analysis chosen for this study to solve Eq. (15)
is the direct integration method because it is suitable for
the nonlinear analysis that is presented in section 4.4.2.
Direct integration is an incremental procedure by which
the displacements, velocities and accelerations are solvedi
at successive time steps At using the displacements, veloc—
ities and accelerations of previous time steps. At each
time step there are 3n unknowns that must be computed.’ They
are q , q and q. The direct integration method used in
this study reduces these to n unknowns by making an assump-
tion about the change in 'acceleration over the time step.
The result is an effective static formulation that can beeasily solved at each step.
4.2·
The base excitation of the structure may be included in
the analysis as known displacements. The finite element ·
equations may be written
Mba Mbb üb Kba Kbb “b Qbg (16)
if ua are the unknown degrees of freedom, ub are known baseexcitations and Qa are known nodal forces. The first set ofequations is
33
Maaüa + Mabüb + Kaaua + Kabub = Qa (17)
Both u and u are displacements in a fixed global refer-a b _
ence frame. Only the displacements in the relative refer-
ence frame, however, result in strains in the system.
Therefore, substituting
“a = q * " (18)
is convenient if n is a vector having the same magnitude as
ua and entries for the horizontal degrees-of-freedom equal
to the base displacements . The vector q is composed of no-
dal displacements in the relative reference frame. Eq.,(l7) _
is rewritten
II=
•· ••i
-Maag+ Kaaq Qa Maan Mabub (Kaan + abüb (19)
A computer program written by the author was used to prove
that·Kaan+ Kabub = 0 (gg)
for the domed structures used. If the mass is also assumed
to be lumped then‘ Mab = 0
(21)· and (19) reduces to ‘
(22)
34
The subscripts 'a' may be dropped since it is‘understood
that q applies to unknown degrees of freedom. One can see
that the effect of the base excitation can be modeled as a
nodal force vector of the mass times the base acceleration.This equation is commonly used in the literature.
The self—weight of the structure is also included in this
study to provide more realistic results. Self—weight is de-
fined as the product of mass and the acceleration due to
gravity. To account for this an acceleration of 32.21
ft./sec.2 is added to each vertical degree of freedom in the
H vector. This results in a constant downward force at
each node which is equivalent to the effect of gravity.
4-3 DAMELEQMHBIXThe damping matrix, C, is unlike the mass and stiffness
matrices in that it is not constructed from element damping
matrices. The damping matrix is only intended to model the
energy dissipative properties of the system. At best, it is
an approximation, however, it should be included to realis- ’
tically model the behavior of a structure.
For Rayleigh damping the damping matrix is approximated
by a proportion of the mass and the stiffness matrices. It.·
can be represented as
C = uM + BK, (23)
35
Values for czand B are determined by solving the two simul-2 taneous equations
By
C! + = =1.,2where—wi
are two frequencies of vibration of the structures
and €Q_are two damping ratios either determined experimen-
tally or assumed.· For structures with many frequencies
Bathe (5) suggests using average values for wiand fi. Engi-
neering judgement may be used to assume these values.
To accurately model damping all of the frequencies of
each structure are considered. Lists of the frequencies ap-
pear in Appendix A and Appendix B. A computer program writ-
ten by Dib (14) was used to find these frequencies. Since
the frequencies of both strucures seemed to be divided into
two groups the average frequency of each group was taken for
wi. A damping ratio of 1% was assumed for the lower wi and
a ratio of 3% was assumed for the higher wi. These E val-
ues were not experimentally determined but were chosen in
order to represent light damping characteristics. For the
small lamella dome the average values for wi and w2 are1.725 hertz and 21.376 hertz, respectively. They yield an a
value of 0.0259 and a B value of 0.00285. For the large la-B
mella dome wl and w2 are 1.608 hertz and 19.198 hertz whichyield an a of 0.0241 and a B of 0.00306.
B
h36
4-4 IHE NEHM8BK;BEIA MEIHQD 24-4-1 Linear Aualxsia
Given the conditions of the nodes at a time t an expres-
sion for the accelerations and velocities at a time t+At can
be derived with the Newmark-Beta method found in Ref. 34.
The resulting equations of this method are
*‘*^'°é3 = (1/ßAt2) <‘”“^"q — tm (25)and
t+^t<ä = tv +· (1/26At) ("+^'°q — ta) (26)
where 9
t .d = tq + tqAt + 1/4 täAt2 (27)
and
- tv = té + 1/2 t&At (28)
Once Eq. (25) and Eq. (26) are derived, linear analysis
by Newmark's method is straightforward. Substituting them
into Eq. (15) yields
(1/BAt2)M(t+Atq—td) + C(tv+(l/2ßAt) <"*^‘*q-"d>> (29) 3
+ K t+Atq = t+AtQ
37
At any time t+At,!b+Atq
are the only unknowns in these
equations. By rearranging the equations the following form
is realized:
· t+At _ t+At^K Q · Q (30) Nwhere
IK = (1/BAt2)M + (1/2BAt)C + K (31)
and
t*^*°é = '°*^*Q + «1/:A·c2>M*‘6 + C((l/2BAt)td - tw G?)Q is referred to as the effective stiffness matrix and in-
cludes the apparent stiffness effects of the mass (inertia
effects) and the damping. A y
4-4)-2 Nsnlinsa; AnalysisIn this section the reduction methods presented in the
preceding section are combined with the Newton-Raphson tech-
nique presented in Chapter 3. The Newmark-Beta method is
used to reduce the dynamic problem, then the Newton—Raphsonl
technique is used to iterate to the solution. The same cri-
terion for convergence as in the static analysis applies to
the dynamic analysis.K
38
If 'k' is the counter of the iteration then the displace-
ment vector
t+At (k) t+At k—l kQ = q( ) + ^q( ) (33)
will be found at each 'k' by solving for 4&q(k) until con-
vergence is reached. The equations of equilibrium for the
nonlinear analysis are
+ + KAq(k) = t+AtQ _ t+AtF(k-l) (34)
At this point Newmark’s method may be implemented. Acceler-
ations and velocities are found by substituting Eq. (33)
into Eq. (25) and Eq. (26). These termsbecome—
k tl) + Aq( ) _ d) (35)
and
respectively. By placing these into Eq. (34) and rearrang-
ing) the following formulation is reached:
^ (k) _ ^ 37KAq - Qeff ( )‘ where
Ö ff = t+AtQ_t+AtF(k-l) _ [(1/BAtz)M (38)e
+ (1/2ßAt)C]t+At§(k-1)-Ctv
39
·c+A1;ä(k-1) __= (t+Atq(k-1)_td)(39)and
12 = (1/BAt2)M + (l/2ßAt)C + 12, <‘*°>
The Newmark method becomes unconditionally stable when the
constant—average-acceleration method, or ß=1/4, is used ac-
cording to Ref. 23.
4-5 9.E TIME §I.E1°§Choice of the correct time step is critical in dynamic
analysis. The time step must be small enough to insure an
accurate solution but not so small that excessive computer
time is needed. Since the Newmark method for B=1/4 is un-
conditionally stable the choice of the time step will be
with regard to the accuracy of the solution. Bathe (4) sug-
gests using a time step At=Tn/10 if TH is the smallest per-
iod of oscillation of the structure to account for the res-
ponse of all modes. For many structures, however, he states
that the primary response is found in the lower modes so
that the higher frequencies and mode shapes may be neglected
and the time steps be lengthened. In these instances the
time steps may be increased to 2;)/10 where TP is the period
of oscillation of the highest frequency considered. Using ·
TP instead of Trlmay increase the time step dramatically and
and still maintain accurate results.
E40
As can be seen in Appendix A and Appendix B the
frequencies of both the small dome and the large dome are
divided into two groups. Only the mode shapes associated
with the lower group of frequencies were assumed to be need-
ed to effectively model the dynamic response. For the small
structure 1.952 cycles per second is chosen as the critical{
frequency because it is the highest of that group. For the
large structure 1.842 cycles per second is chosen. The cor-
responding periods of oscillation, Tp, are 0.512 seconds and
0.543 seconds, respectively. In order to obtain an accurate
solution the time step should be at the most TP/10 or about
0.053 seconds. Also, Wilson, et al. (54) states that for
iearthquake loading there is little justification to use time
steps less than 0.05 sec.‘ The available earthquake data from the El Centro earth-
quake (21) is digitized in increments of 0.020 sec. At a
· time step of 0.020 sec. At/BP would have a value of 0.0377.For Newmark's method this would result in less than 1% per-
iod elongation which is acceptable for many engineering
situations according to Ref. 23. Also, if the time step
used is reduced to 0.020 seconds the earthquake data would
not need to be redigitized which is a complicated process.
.Therefore, for the purpose of convenience and accuracy the‘ time step used in this study is 0.020 seconds.
Chapter V
RESULTS
The WATFIV program that appears in Appendix C was written
by the author for use in this study. It is a finite element
program to model space truss behavior under static load or
dynamic pbase excitation. Nodal displacements and axial
forces in each member can be determined at each load or time
step in addition to nodal velocities and accelerations. The
static analysis capabilities are strictly nonlinearwhilethe
dynamic analysis may be either linear or nonlinear.
Also, Rayleigh damping-may be included as well as a choice
between lumped or consistent mass matrices. The nonlinear
solution technique may be either the Newton—Raphson or the
Modified Newton—Raphson method.
In this chapter results of the program are compared to
results of other researchers to check the program's accura-
cy. Following this, the displacements and stresses caused
by the static loads on both the small and large lamella
domes are presented along with the displacements and maximum
. stresses caused by the live loads and the earthquake excita-
tions. The digitized acceleration data of the El Centro
earthquake of May 18, 1940 was chosen for use in this study
because it was the most severe data available. A comparison
41
42
4is then made of the axial stresses of all members for both
the static and dynamic analyses. The percentage increases
in stress due to the earthquake are then given.
5.1 THE F.LN.I.I.E ELEMENT RBQQBAMui;11 Rrsxious äslilxs
Few sources were found that provided results for a dynam-
ically loaded, nonlinear space truss with which to check the
program. None were found that modeled base excitations.
The simplest comparison found is offered by Nickell (36) who
analyzes the one degree-of-freedom plane truss shown in Fig.
8. In this study, the spring is replaced by a long truss
element that exhibits the same stiffness characteristics.
(The first test performed was a static analysis. A graph of
the load P and the vertical downward displacement of the
joint is shown in Fig. 9. A dynamic test follows where P
increases linearly from 0 to five lbs. in 0.1 sec. and is
held constant thereafter. Both Nickell and the author use a
time step of At=0.1 sec. Results are shown in Fig. 10 and
for both cases the author's results are nearly identical tothose of Nickell.
A larger two dimensional truss system shown in Fig. 11 is
presented by Noor and Peters (38). The horizontal members
have cross-sectional areas of l.6X10’4 m2 while the diagonal
43
P
gi¤7 }„l-—————————g YOU
Ibßn.-—"‘
Figure 8: One degree-of-freedom plane truss presented byNickell.
E44
50 ’°
A- L\11c1<EL L p F "
20
YF .O „~170 _ "
7 2v DISPLACEME/\/T ID.
Figure 9: Displacement vs. load graph for the staticanalysis of the one degree-of-freedom truss‘ presented by Nickell.
45
ISO ^‘/\//CKELI.' ° <>·UL /A/\/A
B25 . 2
$0
E75U4 ’
3.60 · 2Q. '@.26 __
2 2 6 4 6 .7T/ME 666.
Figure 10: Displacement vs. time graph for the dynamicanalysis of the one degree-of-freedom trusspresented by Nickell. «
5u
i46 p
I-~—ll.Om————{————7.0m———•{O.75fT7
P
Figure 11: Ten member plane truss presented by Noor andPeters
· 47
and vertical members have 1.3XlO_4m2 cross-sectional areas.
The Young's modulus of the material is 7.17XlO N/mz and the
density is 2768 kg/m3. These values are the same as those
of aluminum. The load P is an instantaneously applied load
of 4.5X1O4 Newtons acting downward. The critical time step
for stability of the mixed method presented by Noor and Pet-
ers is At=2.63X1O-4 sec. The time step used is O.625XlO”4
sec. Results for both linear and nonlinear analyses are
available.
The displacement of joint A with respect to time for both
linear and nonlinear analyses is presented in Fig. 12. Foru
both cases the author's results match very closely with
those of Noor and Peters. ° Data for the axial force in mem-
ber B with respect to time is also available for linear and
nonlinear analyses. Results are shown in Fig. 13 for the
linear analysis and Fig. 14 for the nonlinear analysis. The
author's results for the linear problem are nearly identical
to those of Noor and Peters. Results for the nonlinear
problem, however, differ slightly. A probable cause is that
the stress-strain curves for the material nonlinearity are
different for both cases. Noor chose to represent the curve
as a polynomial whereas the author chose a bilinear approxi-
mation. A comparison of both appears in Fig. 15. Overall,
the results for the two dimensional truss structure show ex-
cellent agreement with those of Noor and Peters.
48
.60 A-NOOR AND PETERS6- 0114 NA
Q.46LL12 N ONLINEAR ———•··—L5 [,3
TIY„1G;U) .~.75 ‘
__ -4-——————L [ NEÄR
2 4 6 IOTl /\/IEA {T7[[[/.5CC.
Figure 12: Displacement vs. time graphs of joint A forlinear and nonlinear analyses of the ten membertruss.
49A-
/\/OOR AND PETERS0- UL/A/\/A 0>< 704
_2O
/</LL /
·/ xS><Q 5 x
Ü 2 4 68TIMEm2//26600
Figure 13: Axial force in member B for the linear analysisof the ten member truss.
so
xz0" “Z6 ’ ·
• = = = =LL!A [ä (ÜÄOLLL!<c25 „Q A - /\/OORAND PETERS<
08 l.· 0-ULIA./\//X ·
· 2 « 4_ _ 6 8TIME {T7!//IS€C.Figure 14: Aäiää färce igbmemäer B for nonlinear analysis
.511
A- /\/00R AND PETERSo —— U!. /A/\/A
5040
=
A 50
'·<-/)Lu .CKI~co 70.002
.004 .006 .008 .070STRA/N
Figure 15: A comparison of the stress-strain relationships.
i
52
A large three dimensional truss structure is presented by
Noor and Peters (38) and is shown in Fig. 16. The longitu-
dinal members have cross-sectional areas of 0.8X10°4m2, the
diagonals have 0.4Xl0i4 m2 cross-sections and the battens
have 0.6XlO-4m2 cross-sections. The material properties are
the same as those of the previous problem. The time step
used by Noor is At=0.45Xl0_3 sec. and the load is 5000 New-
tons instantaneously applied to the four center nodes indi-
cated in the diagram. Again, the results for the linear and
nonlinear analyses are available.
The displacement of joint A in the direction of the load
for the linear analysis is presented in Fig. 17. The axial
A force in member B as a function of time is graphed in Fig.
18. Both analyses yield nearly identical results. A com-
parison of the displacements for the nonlinear analysis is
shown in Fig. 19 while a comparison of the axial force in
member B is shown in Fig. 20. The results show some devia-
tion of the displacements possibly due to the fact that the
material nonlinearities are different. The graphs exhibit
good agreement in the linear range of the material; however,
when the bilinear assumption of the author applies the dis-
placements deviate slightly.
For all tests the results of the author are in close
agreement with those of other researchers. The only appre-
„ 53 ”
5 m.’? n -.@6 ‘
7.5 m(Ä•·••°fi'
II.
1_$ • ]‘|_§ '“_ l*‘ ?Q? •« 1..: ¤ :1.:
„.Figure16: Large ten-bay truss presented by Noor andPeters.
54
_Ö A -A/OOR AND PETERSE O - UL/LA NA LH7,352
’
A rf ÖLU 2 ...9O \ ,„ \( .~
V) Ü UC3 }4f§ w
0.2 Q2 Q5 0.4 - 06-T//VIE 86C.
Figure 17: The displacement of joint A for the linearanalysis of the ten-bay space truss.
ss
A-NOOR AND PETERS0- UL IANA n
XIO5 0 .AI} //n• 2 0‘ / 1. 1.
Z ., x ·· ..Lö SO [ \ \ LQ: \ \E~'2O6 6.X . \ .<( /1 \ 1. L
x ° ‘0
Q7 Q2 [ 06 0.4 Q5TIME S€C. —
Figure 18: Axial force in member B for linear analysis ofthe ten—bay space truss.
56
7.80‘Q
R7.35 x \E Ö
. <(A’
4. CO
E 45 p A-/x/OOP AND PE TERSo-UL//Ä./\/Ä
Q7 Q2 Q3_ T/ME S€‘C.
Figure 19: Displacement of joint A for nonlinear analysis6 of the ten-bay space truss.
57
.><10$260LL}.U -0- 6- -6 60:
-JE{270 A-/\/OOR AND PET/ERS
O.7 — 0.2 0.5T/ME S€C,
y
Figure 20: Axial force in member B for nonlinear analysisof the ten-bay space truss.
58
ciable differences arrise in the nonlinear problems and maybe traced back to the bilinear assumption of the materialbehavior. _
Before the dome analyses were performed several tests
were run on the large dome in order to check some assump- .
tions that were made. The first test was to determine ifthe use of the Modified Newton-Raphson procedure provides
appreciable savings in computational effort. Two short .static analyses on the large dome showed that the Newton-Raphson procedure required 405 seconds of CPU time and the
Modified procedure needed only 140 seconds. Both analyses
gave identical results. Because of the savings and accuracy
the Modified method was used for the remainder of the test-
ing.
Next, a check to determine the accuracy of Bathe’s sug-gested convergence criterion (See section 3.4) was per-
formed. A static analysis was run using a very small dimen-
sional criterion in. The same test was runusing the nondimensional @=0.001. Both produced identicalresults but the latter used 12% less CPU time. Therefore,
. the criterion that Bathe suggests is used.
59 _
Finally, for the dynamic problem the effect of earthquake
motions in different directions was desired. Two dynamic
analyses on the small dome were performed with the excita-
tions in two perpendicular directions. The results for both F
were identical probably due to the high degree of symmetry
in these domes. Therefore, the direction of excitation does
not affect the results.
5-2 SMLL LAELLA DQME5.2-1 S.t.a1;isAna1y.as.a
Static analyses of the small lamella dome that appears in
Fig. 21 were performed. The nodal load vector Q included
the self—weight of the structure along with the live loads.
The domes are modeled with an additional skin of aluminum
sheet that has a thickness of 0.162 in. in order to closely
approximate the weight of a real dome. With this, the dead
load is about 3 lb./ftkz. In addition, live loads in incre-
ments equivalent to 10 lb./ft.2 or less are added until
failure of the dome occurs.
For a uniformly distributed live load the small dome is
able to withstand 77 lb./ft? plus the dead load before a
singularity in the stiffness matrix is realized. A load-
displacement graph for the downward displacement of node A
and node B appears in Fig. 22. This graph shows the singu-‘
615
— S/NGULAR/TY80 7~00E@-„
60N.ät 61006 @60‘40
Q .6E5Q1 20 · . ‘
. 2 4 6 8‘ DISPLACEMENT en.
Figure 22: Load vs. displacement curves for uniform load onthe small dome.
62
larity occurs without a significant loss of stiffness. The
instability is sudden which would indicate an unstable bi-
furcation. The analysis is stopped after the bifurcationbecause it would constitute a failure of the structure.
Stresses at failure in the numbered members are presented
in Table 1. Only the stresses in a quarter of the structure
are·given due to symmetry. Stresses at various other static
loads are presented in the next section where they are com-
pared to those of the dynamic analyses. None of the members
reached the elastic buckling stress before the global in-
stability. The maximum allowable stresses according to the
Aluminum Association (49) and the buckling stresses for the
numbered elements of Fig. 21 appear in Table 2. ·
A nonuniform load condition was also tested where a full
live load is placed over half of the structure and one half
of this load is placed over the remainder of the structure.
The load-displacement curves for nodes A, B and C are given
in Fig. 23. Node B is in the fully loaded region and node
C is in the partially loaded region. The stiffness matrixof the dome became singular at 67 lb./ft„2 plus the dead
load. Nineteen members reached the buckling stress before
the bifurcation. A complete presentation of the results for
the nonuniform loads will not be given because a dynamic
analysis is not available.
63
TABLE 1Maximum stresses of static analysis for numbered members of
the small dome.
ELEMENT MAXIMUM STRESSNUMBER AT FAILURE (KSI)
1 -5.4112 -5.4483 -5.4214 -17.300 U5 -17.3026 -17.2987 -8.8838 -4.6299 -4.657
10 -8.88111 • -4.65312 -4.63613 -8.88014 -4.64315 27.69816 27.70617 27.706
· 18 27.70119 27.701
64
TABLE 2
Manimum allowable stresses and buckling stresses fornumbered members of the small dome.
‘ (!<SI> <1<s1>”
1 -5.10 -9.89—
2 4 -5.10 -9.89
3 -5.10 -9.89
4 -12.34 -25.43 ‘
5 -12.34 -25.43
6 -12.34·
-25.43
7 -5.10 -9.89 _
8 -4.26 -8.28
9 -4.26 -8.28
10 -5.10 -9.89
11 -4.26 -8.28
12 -4.26 -8.28
13 -5.10 -9.89
14 -4.26 -8.28
15 -12.34 -25.43
16 -12.34 ‘ -25.43
17 -12.34 -25.43
18 -12.34 _ -25.43U
19 -12.34 -25.43
65
80 ‘ QSlNiULAR/TY)/x/00£@ 6
8
Q 60 © NN · Q**“ Q“*— QE ./N ~006@£40Q
2 TOQ4 2
2 46 6 6° DISPLACEMENT }n.
Figure 23: Loacähvs. dä.-splacement curves for uouuniform loadon e sma l dome.
. 66 I
5-2-2 DxuamisAnalyse.;1Dynamic analyses of the small dome subjected to the El
Centro earthquake of May 18, 1940 (21) were performed for
five different load conditions. An accelerogram of the El
Centro earthquake appears in Fig. 24. The load conditions
are:
l. Dead load only.
2. Dead load plus 10 lb./ft.2 live load.3. Dead load plus 20 lb./ft.2 live load.
4. Dead load plus 30 lb./ft.2 live load.
5. Dead load plus 40 lb./ft.2 live load.
Nodal displacements at each time step are calculated inad-·
dition to the maximum axial stresses in each member over the
entire analysis. The analyses are limited to four secondsof earthquake data because the largest ground accelerations
occur within the first four seconds.
In order to include the live load an additional mass isadded to each element that is equivalent to the mass of snow
on the tributary area about each member. By including anacceleration due to gravity (see section 4.2) an effective ·nodal load vector is achieved. The initial nodal displace-
ments of the structure due to this gravity load must also be
included. If they were not then the load would act as if it
were instantaneously applied.
67
03,0.2 I »Ol I ' I · ’ IO O I IIIII
I,, [ I lv ..n n.A.,_lnn. A ’*,~
:; , lvgy 'V ,q, ·. I[I"!1!' ‘, N-0n I __ I I ·f02 ‘ ·_ ,*0-3 4-—y_•o.3zg · — Ground Accelerotion, y
6“ II Ii . I Al.! In lt I,6 ' 7 " V T ' " Y1 ' I-;
-69°•l5.ß8 in/sec Ground Velocity, y
·I6 y,-6.28m. .6
. 4 ”s „s 0
-4Ground Displocement, y
0 6 IO I5 20 25Time, Sec. .
Figure 24: Accelerations, velocities and displacements forthe El Centro earthquake of May 18, 1940.
68
The nodal displacements in the relative reference frame
are calculated since they are the displacements that cause
deformations. The time-displacement graphs of node A in the
direction of the earthquake excitation for all five loads
are presented in Figs. 25 through 29. A composite graph of
the displacements of the four live loads is also given in
Fig. 30. Tables 3 through 7 provide a comparison of the
maximum stresses. For all loads the maximum stresses in the
numbered elements (see Fig. 21) for the static analysis are
compared to those of the dynamic analysis. Stresses in mem-
bers 15 through 19 are not presented for the dynamic analy-
ses since they are the members of the tension ring which are
not included in the fixed base dynamic problem.
5-3 LABQEILQME5.3-1 Ssasisßuzalxsss
The analyses of the large dome are similar to those of
the small dome. For a uniformly distributed live load the
large dome is able to withstand 72 lb./ft.2 plus its dead
weight before a singularity in the stiffness matrix arises.
Load-displacement graphs for the nodes A,"B and C Qf Fig, 31
are given in Fig. 32. As with the small dome, this dome ex-
‘hibits bifurcation instability. The stresses at failure for
the numbered members are presented in Tablel 8. Stresses
69
10-Qoz
-~ t1.0 1.5 -0
-.02 . _
in,.
0t_O _3.0
-.02 _in.
.02
· t3.0 4.0
‘Figure 25: Displacement of joint A for dead load only onthe small dome.
70 I
.02 . °°°
2_O t1.5
-.02tu.
QOZ
t2.0 . 3.0
-.02 '
in..02
° t3_o 4.0
-.02Figure 26: Displacsment of joint A for dead load plus 10
lb./ft. on the small dome.
I
71 g
in..02 - '
· 4 __
-.02in.
.02
-.02 '
in.
'“
-.02Figure 27: Displacezgnent of joint A for dead load plus 20
lb./ ft. on the small dome.
72
in..
.. · t1.0 2'°
-.02 —
in..02
d2.0 -5 3"° t
in..02
3.0 3 .4,0 *5
3 -,02Figure 28: Displacement of joint A for dead load plus 30
lb./ft? on the small dome.
73
in..02
__ __ A _
‘•O2
in.
2 0_5 3.0 T1
-.02V
in.
.3.0 3. 6.0 t
-.02Figure 29: Displacement of joint A for dead load plus 40
lb./ft.2on the small dome.
74
im ..„....... ,. ]_g_]_b_/ ftzz
.......•-- 20 11:./ E1;.2-- —-- so 1b./ ft.2.....----# 40 lb / ft 2 ·/xl
•.• |••‘
••·
·I Ä
•
1.5 2· 1:•• • ••°
~ ·\'/
§\\.,./-Ä
in._md " Q‘~®/geo
, ' _, LO 1:°= y — \ //
in./~.
I^
'
3.0Figure30: A comparison of the displacements of node A for.several liveloads on the small dome.
75
TABLE 3
A comparison of maximum stresses in numbered members of the °small dome for dead lo.ad only.
ELEMENT MAXIMUM STRESSES (KSI)NUMBER STATICDYNAMIC1
-0.1724 -0.49977
27
-0.1724 _ -0.6953
3 -0.1724 -0.7683
4 -0.3775 -1.5300
5 -0.3775 -1.0935
6 -0.3775 -1.3778
7 -0.2248 -0.4164
8 -0.1170 -0.69007
9 -0.1170 -0.6041
10 -0.2248 -0.6106
11 -0.1170 -0.5218
12 -0.1170 -0.7187l
13 -0.2248 -0.6977
14 -0.1170 -0.5987 {15 0.6954 N.A.16 0.6954 N.A.17 0.6954 N.A.
'
18 0.6954 N.A.”18 0.6954 N.A.
I 76 1
1TABLE 4
A comparison of maximum stresses in numbered members of thesmall dome for dead load plus 10 lb./ft. 2.
«‘ ELEMENT. MAx:M¤M sTxEssEs (xs:)NUMBER sTAT:c DYNAMIC
1 -0.9344 -2.05332 _ -0.9349 -2.05053 -0.9347 -2.04704 -2.0470 -4.22465 -2.0470. -4.45166 -2.0470 _ -3.46897 -1.2199 -1.4644
1
8 -0.6346 -1.7334 e
9 -0.6346 -1.733010 -1.2199 -1.464411 -0.6346 -1.734012 -0.6346 -1.711513 -1.2199 E -1.386014 -0.6346 -1.826615 3.7715 ·N.A.
n15 3.7715 N.A.17 3.7715 N.A.18 3.7715 N.A.1 19 3.7715 N.A.
77
TABLE 5A comparison of maximum stresses in numbered members of the
small dome for dead load plus 20 lb./ft. .
NUMBER STATIC ¤YNAM1c1 -1.6696 -2.4123
p 2 e -1.6706 -2.40983 -1.6701 -2.49834 -3.8358 -4.89065 -3.8358-5.13706
-3.8358 -5.21277 6 -2.2403 -2.01458 -1.1650 -2.3430
6 9 -1.1657 -2.2771 618 -2.2395 -2.145311 -1.1658 -2.4295 —12 ” -1.1657 -2.228713 -2.2395 -2.2861 614 -1.1656 -2.339615 6.9397 N.A.
E 16 6.9399 N.A.17 6.9398
9 N.A.18 6.9398 N.A.19 6.9398 N.A.
78
TABLE 6A comparison of maximum stresses in numbered membärs of the
small dome for dead load plus 30 lb./ft. .
ELEMENT MAXIMUM STRBSSBS (xs1> ·NUMBBR STATIC ¤YNAM1c
1 -2.3857 -3.36362 7 _ -2.3874
2-3.3866
3 -2.3865 -3.41724 -5.7541 -6.7501S -5.7539 -6.91136 -5.7541 -6.97797 -3.2925 -2.82838 -1.7134 -3.18441 9 -1.7146 -2.9653
10 -3.2912 -2.931611 -1.7147 -3.2442 E12 -1.7143 -3.022713 -3.2913 -3.073114 -1.7142 -3.155015 10.2090 N.A.18 10.2090 N.A.17 10.2090 N.A.
_ 18 10.2090 N.A.19 10.2090 N.A.
n
79
TABLE 7A comparison of maximum stresses in numbered members of the
small dome for dead load plus 40 lb./ft? .l
NUMBER gTATIc DYNAMIC1 4 «3.l233 -4.33092 I -3.1233 -4.29753 i -3.123 -4.5087
2 4 -7.5331 -9.11665 -7.5331 -9.53516 -7.5331 -9.72537 -4.3104 -3.68368 -2.2447 -4.6477.9 -2.2447 -3.9668
10 . -4.3104 -4.0932ll -2.2447 -4.778612 -2.2447 p -4.069513 -4.3104 -4.356714 -2.2447 -4.492515 13.3654 N.A.15 13.3654 N.A.17 13.3654 N.A.18 13.3654 N.A.19 13.3654 N.A.
81 A
8 . S/NGULAR]7-Y
&NODEN60
ä xx/0066 @ AND ©W- 4 O ' I· Q
WO~¤ 2
2 4 6 8W
DISPLACEMENT m.
Figure 32: Load vs. displacement curves for uniform load onthe large dome.
82 ·
TABLE 8
Maximum stresses of static analysis for numbered members ofthe large dome.
ELEMENT MAXIMUM STRESS ELEMENT MAXIMUM·STRESSNUMBER AT FAILURE(KSI .NUMBER AT FAILURE(KSI)
1 -5.331 21 -2.232E 2 -5.331 22 -5.513
3 -5.333 23 -5.513
4 -11.860 24 -2.2325 -11.860 25 -7.447
6 -11.860 26 -2.2327 -7.370 27 -5.513
8 -3.484 28 -5.513
9 -3.484 29 -2.232
10 -7.370 30 -7.439
11 -3.484 31 -2.23212 -3.484 32 -5.511
13, -7.370 33 32.355
14 _ -3.434 34 32.603
15 -9.654 35 32.357
16 -9.657 36 32.356
17 -9.655 37 32.60218 -9.655 38 32.35719 -9.655 39 32.357
%20 -7.447 40 32.603
n l
834
for other loads are given in the next section where they arecompared to the dynamic results. The maximum allowablestresses and the buckling stresses for the numbered elements
of Fig. 31 are given in Table 9. Again, none of the members
·reached the buckling stress before the global instability.
For the nonuniform load the stiffness matrix becomes sin-
gular at 45 lb./ft.2 plus the dead load. Fig. 33 shows the
load-displacement graph for the downward deflections ofnodes A,B,C,D and E. Nodes D and E are in the lesser loaded
portion of the dome. In this dome, unlike in the smalldome, no members reached their buckling stress prior to glo-
·bal instability. I
5.3.2 Dynamic AnalyacaDue to the cost of analysis of the large dome the loading
conditions are reduced to three. They are:
1. Dead load only.
2. Dead load plus 20 lb./ft? live load.3. Dead load plus 40 lb./ft? live load.
- Nodal displacements at each time step are calculated in ad-
dition to the maximum stresses in each member.
The time displacement graphs of node A in the directionof the earthquake excitation for all three load conditions
are shown in Figs. 34 through 36. A composite graph of the
84
TABLE 9Maximum allowable stresses and buckling stresses for
numbered members of the large dome.
‘ (xs:) „ < KSI)1 -5.10 g -9.892 -5.10 -9.89
3 -5.10 -9.89
4 -12.34 -25.43
5 -12.34 -25.43 g6 -12.34
4-25.43
7 -5.10 -9.89
8 -4.26 -8.28
9 4 -4.26 -8.284
10 -5.10 -9.89
11 -4.26 4 -8.28
12 -4.26 -8.28
13 -5.10 -9.89
14 4-4.26 -8.28
15 -12.34 -25.43
16 -12.34 E -25.43
17 -12.34 -25.43
18 -12.34 -25.434
19 -12.34 -25.43
85
TABLE 9 (cont ' d)
ELEMENT MAXIMUM ALLOWABLL BUCKLINGNUMBER COMPREQSIVE STRESS STRESS ‘
20 -5.10 -9.89‘2l -4.65 -9.03
22 -4.65 -9.03_ 23 -4.65 -9.03
24 -4.65 -9.0325 -5.10 -9.8926 -4.65 -9.0327
3-4.65 -9.03
28 -4.65 -9.0329 -4.65 -9.0338 -5.10 -9.8931 -4.65 -9.0332 _ -4.65 -9.0333 -12.34 -25.433* -12.34 -25.4336 -12.34 -25.4336 -12.34 -25.4337 -12.34 -25.4336 -12.34 -25.4330 -12.34 -25.43*0 -12.34 -25.43 ”
86I
50 . gNODE7v001s©406
SSO
2 7Q<L'O·~*7O
7.0 62.0 :20 4.0. 6.0DISPLACEMENT m.
Figure 33: Load-displacement curves for nonuniform load onthe large dome.
I
87
in.-.02
6 6 2-0.+.:.
-.02in.
QOZ
-.02in.
002
3.0 4 3.5 6.0 ”°
'°•O2 AFigure 34: Displacement of joint A for dead load only on
the large dome.
F F88
1:1.OO1
1..0 1_ 2.0 C
*0•1
in. „0.1
2.0 3.0 t
-0.1
in.0•1
3.0 6.0 t F
-0.1 ‘
Figure 35: Displagéement of joint A for dead load plus 20lb./ft. on the large dome.
. 89A
111. '0.1
2.0 6
-0.1in.
l ti2,0 2.5 .0
TO-1
in.0-1
*
·
, 3.5 _ 6.0 t
—0O1
Figure 36: Displacäement of joint A for dead load plus 40„ lb./ft. on the large dome.
90
two live loads is presented in Fig. 37. Tables 10 through
12 offer comparisons between the maximum stresses in the
numbered elements of both the static and dynamic analyses.
The tension ring elements are not included in the dynamic
analyses.
5-4 9.EEQRBQIHL.é.B§EEJ@§[email protected].$The allowable stress for all members according to the
Aluminum Association (49) were calculated. Allowable ten-
sile stress for all members is 19 ksi. Allowable compres-
sive stresses are found in Table 2 for the small dome and in
Table 9 for the large dome. These compressive stresses are
very low because the members are unusually long for their
cross-sectional areas.
For the static analyses all members of the small dome
were over-stressed at failure. These over-stresses ranged
from 6% in members 1, 2 and 3 to 74% in members 7, 10 and
13. No members, however, reached the buckling load or yield
stress prior to failure. For the large dome, half of the
members were over-stressed at failure. The over—stresses”
ranged from 5% in members 1, 2 and 3 to 71% in members of
the tension ring. Other over-stressed members were 7, 10,
' 13, 20, 25, 30, 22, 23, 27, 728 and 32. As with the small
dome, no members reached their buckling stress and none
reached the yield stress of the material.
91A in. -———— 20 1.b,/ ft.2
.. - 40 110./
1.5tn,0.0
0
in.
0.-0 0Figure 37: A comparison of the displacements of joint A for
the two live loads on the large dome.
I
92
TABLE 10A comparison of maximum stresses in numbered members of the
large dome for dead load only.
ELEMENT MAXIMUM sTREssEs (xs:)NUMBER sTAT1c 0xNAM1c
1 -0.1786 -0.61092 “ -0.1787 -0.64753 -0.1787 -0.58354 -0.3594 -0.91855 -0.3594 -0.96196 -0.3594 -0.89957 -0.2235 -0.47558 -0.1159 -0.47759 -0.1159 -0.5595
18 -0.2236 -0.465111 -0.1158 -0.4421
E 12 -0.1156 p -0.358213 -0.2236 -0.505914 -0.1158 -0.464315 -0.2558 -0.626415 -0.2558 -0.628417 -0.2558 -0.746218 -0.2558 -0.737318 -0.2558 -0.8282
93 !
TABLE 10 (cont'd)
¤·¤¤¤~¤·NUMBER STATIC DYNAMIC20 -0.2264 ° -0.3024
21 -0.0622 _ -0.473522 -0.1722 -0.477623 -0.1721 -0.471324 -0.0622 -0.417725 -0.2264 -0.290926 -0.0624 -0.346027 -0.1722 -0.5168
28 -0.1722 • -0.490829 -0.0624 -0.3659 230 -0.2262 -0.2864 „
31 -0.0623 -0.3277
32 -0.1722 -0.514733- 0.9694 N.A.
34 0.9808 N.A.
35 0.9694 N.A.
36 0.9694 N.A.
37 0.9808 N.A.
38 0.9694 N.A.39 0.9694 N.A.40 0.9808 N.A. l
94
TABLE llA comparison of maximum stresses in numbered membärs of the
large dome for dead load plus 20 lb./ft. .
— MAx1M¤M STRBSSES (xs:)STATIQ DYNAMIC
1 #1.550-6.43232
' -1.550 -6.43593 -1.550. -6.1703
44 -3.154 -8.30035 -3.154 -9.8866
h 6 -3.154 -8.9202. 7 -1.959 -5.8969
8 -1.008 -3.60169 -1.008 -4.2084
10 -1.959 -5.191511 -1.008-3.528012
-1.0081
-4.033613 -1.959 -5.953114 -1.008 -5.058315 -2.263 -7.6324 U16 -2.263 -7.569117 -2.263 -7.723718 -2.263 -7.150718 -2.263 -7.2858
95 _
TABLE ll (c¤¤t'd)
EEEEEEENUMBER_ STATIC DYNAMIÜ
,20 -l.983.' -3.5416
21 -0.547V V-3.8718
22 -1.508 -4.101023 -1.508 -4.3558
E 24 -0.547 -3.3052
25 -1.983 -3.194526 -0.547 -3.353027 -1.508 -3.8117
28 -1.508 -3.9527 'V
29 -0.547 -3.1650
gg -1.983 -2.7943 .31 -0.547 -2.950732 -1.508 -4.8165
33 V8.506 N.A.
34 ·8.506 V N.A.
35 8.506 N.A.
36 8.506 N.A.37 8.506 N.A.
38 8.506 7 N.A.
39 ·8.506 N.A.40 J 8.506 N.A.
V
l96
TABLE 12A comparison of maximum stresses in numbered memgers of the
large dome for dead load plus 40 lb./ft. .
1 NVMBER s·1·A·r1c: owmxxc ‘
1 -3.021 -8.8923
2 -3.023 -8.7624
3 -3.022 -9.1952
4 -6.430 -13.1460_ 5 -5_43Q -14.6910
· 6 _6•43O -12.41407 -3.973 -8.14688 -1.981 -4.33689 _l_98l -4.5931
10 _3_975l
-8.165711 -1.980 -5.1791 .12 _l_98O -4.435913 _3_975
c -8.712214 _l•98O ‘ -7.001015 -4.895 -8.610216 -4.895 ' -8.6628
‘ 2
17 -4.894 -9.691918 -4.894 -11.647019 -4.894 -13.4650
I
97
• TABLE 12 (cont'd)
RRR20 -4.033 -5.3766
21 -1.149 -5.2154 622 -3.037 -6.3230
23 -3.035 -6.4656
24 -1.149 “ -4.6812
25 -4.033 -4.0493Z6 -1.153 -4.397627 -3.063 -6.5516
28 -3.036 -6-3407
29 -1.119 -5-6663
30 -4.029 -8-7535R
31 -l_l5l -7.4687
32 -3.036 -8-8381
_ 33 17.387 8-6-34 17.520 8-6-
35 17.388 8-6-36 17.388 ‘ 8-6-
. 37 17.569 _ N.A.
38 17.388 8-6-38 17.366 8-6-88 17.570 8-6-
98
In the small dome no members became over-stressed up toand including the static load of 40 lb./ft.2 . The highest
stress was 6% under allowable. For the large dome , also,
no members were over-stressed up to and including 40
lb./ft.2. The highest stress was 8.5% under allowable for
the tension ring members.V
For the dynamic analyses significant increases in maximum
stresses were noticed for all loads on both domes over the
stresses in the static analyses. Members of the small dome
experienced average increases in maximum stresses of 303%,
121%, 51%, 40% and 47% for the dead load only, dead loadplus 10 lb./ft.2, dead load plus 20 lb./ft.2, dead load plus
30 lb./ft.2 and dead load plus 40 lb./ft.2 , respectively.
The maximum increases in stress of any member for the loadsgiven were 514%, 187%, 108%, 84%, and 113%, respectively.
The stresses throughout the analyses remained within the al-
lowable range for all cases except 40 lb./ft.2 where members
8, 11 and 14 were 8.5% over-stressed but still remained be-
low the buckling stress.
Members of the large dome experienced average stress in-
creases of 236%, 261% and 191% for dead load only, dead load
plus 20 lb./ft.2 and dead load plus 40 lb./ft.2, while themaximum increases in stress for any member were 661%, 608%
and 549%, respectively. These increases resulted in many
99 1
over-stressed members and even some buckled members. For 20
lb./ft. members 1, 2, 3, 7, 10 ,13, 14 and 32 were over-
stressed to a maximum of 26% over allowable but all below
the buckling stress. For 40 lb./ft.2 all members were
over—stressed except members 16, 17, 18, 25 and 26 to a max-
imum of 90%. Ten elements reached buckling, however,all of
these elements were outside the quarter region of Fig. 31.
A final note on the results should be given. At no time
during the dynamic analyses did either dome become globally
unstable. All analyses were able to be carried to comple-
tion. This may be because the entries of the force vectordue to the earthquake excitation act in a plane parallel to
the ground. For shallow domes like the ones used in this
study horizontal forces are probably less likely to cause a
global instability than vertical downward forces.
Chapter VI
CONCLUSICNS AND RECCMMENDATIONS
6-1The question to be answered by this study is by how much
does an earthquake excitation increase the axial stresses ina reticulated space truss such as the domes in Figs. 2 and
3. The stress increases due to the excitation were signifi-
cant especially in the large dome where the average increase
was about 200% over the stresses of the static analyses.
For the small dome the stress increases were smaller but
still over 40%. Earthquake excitations cause stress in-creases that must be considered in the design of a dome in a
seismic region.
Even though these stress increases appear to be large,
these domes exhibit surprising strength characteristics un-
der dynamic loading. Many studies have shown that these
domes are strong for their weight when subjected to static
loads and for dynamic loads they again appear to be strong.
For example, the large dome was able to withstand a distri-
buted load of nearly six times its own weight before its
members became over-stressed under the base excitations.
The small dome was able to withstand an additional load of
nearly 13 times its own weight. This is impressive and it
100
101
demonstrates the strength and usefulness of these domes in
seismic regions. Overall, the lamella dome is a strong
structure and able to resist base excitations based on the
results of this study.
6.2 BE§QMENQAIlQN§ EQB EHIHBE‘EE§EABQHAlthough the basic questions of this study were resolved,
many other questions were raised during the research that
were not answered due to the constraints on time. Several
extensions of the thesis topic that address these questions
are given. Other students may wish to pursue these topics.The computer program developed for use in this study may
be refined to make it more versatile and economical. It may
be modified to include various types of dynamic loads such
as blast loads or wind loads to provide more complete dynam-
ic analysis capabilities. The program may be refined to in-
clude different material nonlinearities so that different
materials may be considered and more accurate results pro-
duced. Other direct integration methods may also be includ-
ed. The Wilson method or the Houbolt method might be suita-
ble for other structures. A study to find the most
economical direct integration scheme for these domes is also
a possible extension. .
. 102 A
The most time consuming step in a typical computer analy-sis using this program is arranging and inputting the data.
The program by Dib (14) produces the necessary data for a
lamella dome or a geodesic dome and numbers the elements and4
joints to provide a reduced bandwidth. The two programs may _be combined for a more efficient program.
Additional testing may be performed to answer some in-
triguing questions. A geodesic dome or other popular dome
configuration with the same weight and overall dimensions as
the lamella could be tested with the same earthquake data.The results may be compared to the results of this study.
Also, the nonuniform loads mentioned in section 5.2.1 may be
included and dynamic analyses performed. A comparison with
uniform loads would be possible. The methods presented by
Rosen and Schmit (47) and (48) to model imperfect members
could be included in the program to study the effects of
member imperfections.n
The frequencies of the structure are calculated using the
stiffness matrix of the undeformed structure and the mass
matrix of the unloaded structure. By increasing the mass
with the live load and changing the stiffness matrix with
the initial displacements caused by the live loads, the fre-
quencies should change. A study could be made on how the
frequencies change with different live loads. Any changes
103 I
in the frequencies would alter the damping coefficients andtime steps. The amount of change in these parameters should
I
be determined to see if they are significant and to what de-
gree they affect the analysis. If the time steps would vary
significantly for different loads then it may be advanta-
geous to generate an artificial accelerogram to accept the
varying time steps.
The last question to be answered is whether the bending
strains should be included in the analysis. Perhaps the as-
sumption of the space truss element is not accurate. In-
cluding the live loads may require that a space frame ele-ment be used. The program may be modified to model a space
frame element and comparisons between the frame and thetruss can be made.
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Appendix A
FREQUENCIES OF VIBRATION FOR THE SMALL LAMELLADOME
FREQUENCY FREQUENCY FREQUENCY FREQUENCYNUMBER (HERTZ) NUMBER (HERTZ1
1. 1.3993 18. 14.6974
2. 1.4014 19. 17.9737
3. 1.7227 20. 17.9742
4. 1.7290 21. 19.8414
5. 1.7523 22. 19.9268
6. 1.7764 23. 19.9268
7. 1.7995 · 24. 20.6390
8. 1.8205 25. 22.2076
9. 1.8999 26. 22.2037
10. 1.9515 27. 25.3070
11. 2.4822 28. 25.3088
12. 7.6420 29. 25.7914
13. 9.4609 30. 25.7914
14. 9.4612 31. 28.0129
15. 14.4240 32. 28.013316. 14.4240 33. 28.999117. 14.6974
11O
E
a
’ Appendix B
FREQUENCIES OF VIBRATION FOR THE LARGE LAMELLADOME
FREQUENCY FREQUENCY FREQUENCY FREQUENCYNUMBER (111:12*1*2) NUMBBR <H12:R·rz>
1. 1.345 22. 1.752
2. 1.345 23. 4 1.754
3. 1.359 24. 1.798
4. 1.359 25. 1.798
5. 1.364 26. 1.8286. 1.367 27. 1.828
7. 1.367 28. 1.841 ·8. 1.395 29. 1.8419. 1.395 30. 1.842
10. 1.417 31. 2.687
11.l
1.417 32. 4.906
12. 1.423 33. 6.560
13 1.479 34. 6.560
14. 1.730 35. 8.360U15. 1.735 36. 8.360
16. 1.735 37. 8.43017. 4 1.742 38. 9.56318. 1.742 39. 9.563
19. 1.746 40. 11.636
n 20. 1.746 41. 11.63621. 1.752 42. 12.384
111
112
Appendix B (Continued) 4
FREQUENCY FREQUENCY FREQUENCY FREQUENCYNUMBER (HERTZ) NUMBER (HERTZ)
43' 12.384 64._l
17.9064‘* 12.442 65. 18.0624S' 12.443 66.
518.062
46' 12.443 67. 18.89247' 13.412 68. 18.89248* 13.412 69. ,19.66649° 13.775 70. 20.851SO' 14.372 71. 20.851Sl' 14.372 72. 21.500S2' 15.038 73: ‘ 21.625 _58* 15.609 74. 21.625S4' 15.609 75. 21.94455* 16.891 76. 21.944
5 S6' 17.071 77. 22.37057* 17.071 78. 22.37058* 17.359 79. 22.762S9' 17.359 80. 22.76260' 17.437 81. 23.3916l° 17.437 ' 82. 23.80462° 17.540 83. 23.80463° . 17.540 84. 24.337
113
Appendix B (Contiuued)
FREQUBNY FREQUENCY FREOUENCY FREQUENCYNUMBER (HERTZ) NUMBER (HERTZ)
°85. 24.337 90. 25.592
86. 24.792 91: 25.592
87. 24.792 92. 25.858
88. 25.285 93. 26.198
89. 25.285
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jTHE EFFECTS OF EARTHQUAKE EXCITATIONS
ON RETICULATED DOMES
David A. Uliana(ABSTRACT)
Comparisons were made on the behavior of two full-sized
reticulated domes subjected to uniform static loads only and
uniform static loads with earthquake excitations. Space
truss elements were used in the dome models. The stiffnesse
matrix of the space truss element allows for the nonlinear
strain-displacement behavior and the stress-strain behavior
of the material is modeled with a bilinear approximation.
The nonlinear solution technique is the Newton-Raphson meth-
~ od while the direct integration technique is the Newmark-
Beta method.The joint displacements for the static and the dynamic
analyses were compared for both domes along with the axial
stresses in all members. The percentage increases in the
”axial stresses of the dynamic analyses as compared to those
of the static analyses were determined.
The reticulated domes used in the study were found to betcapable of withstanding the earthquake excitations when sub-
jected to various uniform loads without failure. ·