AD-780 425Rotating Annular Disk 23 Modulus Variation for Rotating Orthotropic Disk with no Edge...
Transcript of AD-780 425Rotating Annular Disk 23 Modulus Variation for Rotating Orthotropic Disk with no Edge...
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COMPOSITE DESIGN SYNTHESIS
BATTELLE COLUMBUS LABORATORIES
PREPARED FOR
OFFICE OF NAVAL RESEARCH
ADVANCED RESEARCH PROJECTS AGENCY
JUNE 1974
DISTRIBUTED BY:
AD-780 425
KHr National Technical Information Service U. S. DEPARTMENT OF COMMERCE
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Und ass i f ird SECUHITY CLASSIFICATION Of-' THIS PAC.t Clthiri DIIIII /irli-fi-.-fJ /^z) z?0 f^:
REPORT DOCUMENTATION PAGE I. REPORT NUMÜf-R 2. GOVT ACCLSSION NO
RF.AD INSTRUCTIONS BEFORE COMPLETING FORM
3 RECIPItNT'S C AT ALOG NUMBLH
4. TITLE fand Suhtlllr)
Composite Design Synthesis
7. AUTHORf.v>
Milton Vagins
9. PERFORMING ORGANIZATION NAME AND ADDRESS
Bat'.ielle's Columbus Laboratories 505 King Avenue
Columbus, Ohio 43201
5 TYPE OF REPORT A PERIOD COVERED
Final Techni:al Report, 1 year
6 PERFORMING ORC. REPORT NUMbER
6. CONTRACT OR GRANT NUMBER^)
N0001/4-72-C-0195
II. CONTROLLING OFFICE NAME AND ADDRESS
Office of Naval Research 800 North Quincy Street Arlington, Virginia 22217
T« MONITORING AGENCY NAME * ADDRESSfi/ di»efpm from Conlrolllnt Ollice)
Advanced Research Projects Agency 1400 Wilson Boulevard Arlington, Virginia 22209
10. PROGRAM ELEMENT. PROJECT, TASK AREA & WORK UNIT NUMBERS
12. REPORT DATE
June 14, 1974 13. NUMBER OF PAGES
80 15. SECURITY CLASS, (ol Ihlt report)
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Approved for public release; distribution unlimited.
17. DISTRIBUTION STATEMENT (of the abilract enltied In Block 10, II dlllerrnl Iron Report)
18. SUPPLEMENTARY NOTES
19. KEY WORDS (Continue on re"ef«e «id« 11 neceefry mnd Identify by block number)
Composite Design Synthesis Rotating Disks Pressurized Thick-wall Cylinders
NATIONAL TECHNICAL INrORMATlON SERVICE
U 8 Department of Commerce Springfield VA 22151
20. ABSTRACT (Continue on reverie »id« (f noceseery end Identity by block number)
The research carried out deals with the problem of design synthesis in heterogeneous elasticity. Design synthesis is defined as the achievement of a des red design criterion, i.e., stress distribution, strength-to-weight ratio, etc., by preselecting a stress or displacement pattern in a stretched plate and then determining the variation of the elastic moduli that is required to permit the desired effects. This accomplishment requires the solution of the governing equations of %o
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elasticity, particularly the compatibility equation, in terms of preselected stress fields in the body of the plate for unknown material \. -operties which are spatial functions.
During this work, solutions to the moduli variation problem for annular disks have been achieved for two stress criteria: constant hoop stress, and constant in-plane shear stress. The disk may be rotating and have boundary traction. A computer program, D0M0V1, was developed to carry out these solutions. Attempts to solve the moduli variation problem for a hole in an infinite plate (two-dimensional), subject to certain stress distributions, were unsuccessful. However, great insight was gained into this problem for future work.
Basically, this work has shown that the concept of design synthesis, as defined here, is a workable discipline and,in the case of rotating annular disks and pressurized thick-wall cylinders, can be applied utilizing the present state-of-the-art fabrication technology, but that its application to complex problems requires additional work.
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Unclassified SFClimTY CLASMTIC ATIOM OF THIS PAGF'Whon IlKIn hnli-tr,lt
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FINAL TECHNICAL REPORT
COMPOSITE DESIGN SYNTHESIS
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OFFICE OF NAVAL RESEARCH DEPARTMENT OF THE NAVY
June 1A, 1974
Milton Vagins
D D C prapmni?
UüisisEüins c
BATTELLE Columbus Laboratories
505 King Avenue Columbus, Ohio 43201
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Battelle is not engaged in research for advertising, sales promotion, or publicity purposes, and this report may not be produced in full or in part for such purposes.
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June 14, 1974
llBäireiie Columbus Laboratories 505 King Avenue Columbus, Ohio 43201 Telephone (614) 299-3151 Telex 24-5454
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II fl II 0 n r r o 0 ö
Director Advanced Research Projects
Agency 1400 Wilson Boulevard Arlington, Virginia 22209
Attention Program Management DODAAD Code HX 1241
Dear Sir:
Form Approved Budget Bureau No. 22-R0293 Contract No. N00014-72-C0195
Enclosed is the Final Technical Report on the cc "-ract titled "Composite Design Synthesis".
Sincerely,
Milton Vagins Principal Investigator Applied Solid Mechanics Section
MV:Jlg
Enclosures (3)
iV
T "H-
Final Technical Report
on
Composite Design Tynthesis
14 June 1972
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AR PA Order Number
1977
ProKram Cod« 1 Number
2N10
Contract Number
N00014-72. •C-0195
Scientific Officer
Director Structural Mechanics, Mathematical and Information Science Division
Office of Naval Research Department of the Navy
Name of Contractor
Batteile Memorial Institute Columbus Laboratories
505 King Avenue Columbus, Ohio 43201
Principal Investigator
Milton Vagins Phone (614) 299-3151 Ext. 1659
Effective Date of Contract
30 December 1971
Contract Expiration Date
28 February 1974
Amount of Contract
$39,400.00
Sponsored by Advanced Research Projects Agency
ARPA Order No. 1977 dated 18 October 1971
The views and conclusions contained in this document are those of the author and should not be interpreted as necessarily representing the official policies, either expressed or implied, of the Advanced Research Projects Agency or the U. S. government.
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TABLE OF CONTENTS
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REPORT SUMMARY i
SYMBOLS ii
INTRODUCTION 1
GENERAL DISCUSSION 4
APPLICATIONS 12
Rotationally Symmetric Problems 12
Non-Syinmetric Problems 35
Limits and Other Considerations 42
Two-Diinenslonal Boundary Considerationr ^';
Material Considerations 48
SUMMARY 49
REFERENCES 50
APPENDIX A
INPUT DATA FOR PROGRAM D0M0V1 A-l
APPENDIX B
LISTING OF PROGRAM DOMOVi B-l
APPENDIX C
OUTPUT OF PROGRAM DOMOVI C-l
LIST OF FIGURES
Figure 1. Pressurized Annular Disk 13
Figure 2. Modulus Variation for Orthotropic Disk, with Pressur- ized I.D., with aa ■ Constant 18
Figure 3. Modulus Variation for Isotropie Disk with Pressurised I.D., with aQ = Constant 19
Ml
x
■ i * Ifr i^
^
LIST OF FIGURES (Continued)
Pag(
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Figure 4.
Figure 5.
Figure 6.
Figure 7.
Figure 8.
Figure 9.
Figure 10.
Figure 11.
Figure 12,
Figure 13
Figure 14
Modulus Variation for Orthotropic Disk with Presour- ized O.D., with 0o = Constant 21
0 Modulus Variation for Isotropie Disk with Pressur- ized O.D., with o ■ Constant 22
Rotating Annular Disk 23
Modulus Variation for Rotating Orthotropic Disk with no Edge Loads, with o - Constant 29
Modulus Variation for Rotating Orthotropic Disk, Ortho- tropic Ratio of 2.5 with no Edge Loads and a0 = Constant 30
Modulus Variation for Rotating Orthotropic Disk with no Edge -oads, and with Constant In-plane Shear Stress . . 31
Modulus Variation for Rotating Orthotropic Disk Ortho- tropic Ratio of 2.5 with no Edge Loads and with Constant In-plane Shear Stress 32
Modulus Variation for Orthotropic Rotating Disk Stressed on O.D. RPM - 2000; Density =0.1 lb/in.3; Stress on O.D. ■ 10,000 Psi; O.D. ■ 60.0 inches; I.D. ■ 6.0 inches, aQ ■ Constant
Mo ulus Variation for Orthotropic Rotating Disk Stressed on O.D.: RPM - 4000; Density - 0.1 lb/in.3; Stress on O.D. - 12,000 Psi; O.D. - 60.0 inches; I.D. ■ 6.0 inches, a« - Constant 34
Small Hola in Infinite Plate Subjected to Uniform, Uniaxial Tensile Stress Field 36
, Two-Dimensional Finite-Difference Model 40
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REPORT SUMMARY
This is the Final Technical Report describing the work done and
the accomplishments of the research effort for ONR-ARPA titled "Composite
Design Synthesis".
The research carried out deals with the problem of design syn-
thesis in heterogeneous elasticity. Design synthesis is defined as the
achievement of a desired design criterion, i.e., stress distribution,
strength-to-weight ratio, etc., by preselecting a stress or displacement
pattern in a stretched plate and then determining the variation of the
elastic moduli that is required to permit the desired effects. This accom-
plis'iment requires the solution of the governing equations of elasticity,
particularly the compatibility equation, in terms of preselected stress
fields in the body of the plate for unknown material properties which are
spatial functions.
During this work, solutions to the moduli variation problem for
annular disks have been achieved for two stress criteria; constant hoop
stress; and constant in-plane shear stress. The disk may be rotating and
have boundary traction. A computer program, D0M0V1, was developed to carry
out these solutions. Attempts to solve the moduli variation problem for
a hole in an infinite plate (two dimensional), subject to certain stress
distributions wire unsuccessful. However, great insight was gained into
this problem for future work.
Basically, this work has shown that the concept of design syn-
thesis, as defined here. Is a workable discipline and in the case of
rotating annular disks and pressurized thick--'all cylinders, can be
applied utilizing the present state-of-the-art fabrication technology,
but that its application to complex problems requires additional work.
VJIII
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SYMLOLS
Symbol Quantity
a,b inside and outside radius
a. material coefficients
il
[1
c dimensionless shear-modulus coefficient ! ='^7,)
A .B »G .C,.C„ constants of integration o' o' o' 1' 2
orthotropic ratio [-p) e
E. modulus of elasticity corresponding to the subscripted direction (i • r,9)
';|] i A f1(e),f2(9) functions of 9 only
G modulus of rigidity ( = G Q = 1 00
g acceleration of gravity
k radius ratio ( =b/a)
\ i-t k. orthotropic ratio ( =e)
k- Poisson's ratio in tangential direction ( svfl )
n cos 0
n sin 0
P internal pressure
q external pressure
r radius
T temperature difference function
u,v displacements
R,X,Y body forces
U specific strain energy
U total strain energy
•■■;• II
^
I n
(I i
I'
l ii
U Q 0 II 0 e
i
Symbol
or
0
Y
e
I
9
\
v
Y
U)
6
n
Quantity
coefficient of linear thermal expansion
exponent (2-k,/k«)
material density
strain component wich one or two subscripts
body force in tangential direction
angular position
exponent
Poisson's ratio in tangential direction ( - vQ ) ü r
exponent ( H - [(J , ^ ]) -1C1 ■ K^ ^>
X/J ) J\r\m
dimensionless position ratio (=r/b)
coordinate system offset angle
stress function
rotational velocity
variation factor
arbitrary function
*.
J 11
■^r wm mm^^i ^
K. - — —-
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n fi n
11
i.
INTRODUCTION
It has long been recognized that structural ele.nents composed of
composite materials, such as glass, boron, carbon, or other filaments, embedded
in a suitable matrix, such as epoxy or polyester, offer outstanding strength-
to-weight ratios. The potential of such materials is considered so great that
material scientists and enginaers believe that they will form the bulk of the
structural materials of the future. Though the application of such materials
has been a growing part of the state of the art for structural components,
particularly in the aerospace industry, thi translation of the concept of
fibrous composites into a primary load carrying structure has been and remains
a challenging process.
The present and growing u^e cf structural elements fabricated from
composite materials creates the need for the development of a rational analytic
design basis, which to a great extent is presently non-existent. This is not
to be construed as meaning that little or no rtcearch on composite materials
has been carried out. On the contrary, a large amount of literature has been
generated dealing with both the determination of the mechanical properties of
these materials and the analysis of specific structures fabricated from them.
In general, from the microscopic viewpoint, research on the mechani-
cal properties of composites has dealt with the determination of such properties
for materials having given component elements ordered in fixed spatial
relationships. The spatial relation in these cases might have a high degree
of symmetry, as in long or continuous filament composites, or a completely
random or homogeneously disordered array as usually employed in short carbon
or boron fiber composites. Such research has been directed towards the
creation of analytic or experimental methods of determining the mechanical
properties of composite structures in terms of the known properties of the
composites' components, characteristic of »"h^ approach are the works of
Sayers and Hanley [1] , Chen and Cheng [2], Hill [3], and Gaonkar [4], among
others.
In the analysis of structures composed of such composites, the
material has generally been treated as exhibiting gross, homogeneous.
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Numbers in brackets are references found at the end of this report.
' \ mm
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(I Isotropie or anisotropic mechanical properties. These gross properties, when
entered into the constitutive equations defining the material, have allowed
analyses of such st uctures through the classical methods of the theory of 1
elasticity, plates and shells, vibration, and others. Along these lines, the \
concept of the "unidirectior.il lamina" was introduced and utilized as the l
"fundamental unit of material" in design and analysis. This procedure I
employs test data obtained from a unidirectional lamina as the basis for the
design of laminated components and structures. Exemplary work done along
these lines has been carried out by Dong [5], Tsai [6,71, Tsai and Azzi [8],
Whitney [9], and Whitney and Leissa [10,11], also among others.
All of these analyses have dealt with materials and structures that
have predetermined mechanical and behavioral characteristics. That is, once
the geometric array and the constituents of the composite are prescribed,
then the mechanical properties of the material and the response characteristics
of a structure fabricated from such a material have been inherently established.
An;lysis will merely determine what these properties and response charac-
Ceristics are.
When dealing with composite materials, the analytical procedures
discussed above appear to be highly inefficient in many applications. The
designd. of fibrous composite structures is presented with numerous degrees
of freedom and an opportunity to exercise ingenuity totally unavailable to him
with conventional materials. Composites, whether filanentary, fibrous, or
sintered or fused metallics, are capable of being tailored to meet specific
requirements. When considering specific structural applications for such
materials, it would be logical to assume that a structure could be optimized,
depending, of course, upon the optimization criterion, by varying the mechani-
cal properties of the material throughout the structure. Further it would be
logical to bypass analysis completely and define this now nonhomogeneous
structure by some means of design synthesis. Admittedly, the creation of a
design synthesis procedure to adequately handle most problems in structural
design is quite difficult. However, the concept of design synthesis tc deter-
mine the variation of the mechanical properties of a material within a structure
so as to achieve a desired stress or deformation pattern ir. that stvueture is n
one capable of being developed.
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The work described herein deals with the first pi.ase of an effort
to develop a design synthesis methodology for composites. It is directed
specifically to the problem of heterogeneous plane elasticity.
(1 i II II
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GENERAL DISCUSSION
Consider the following question:
Given a plane elastic body with known boundary tractions and/or
displacemtnts, can the mechanica) properties of the continuum be described
such that an "arbitrary" stress distribution within the body is met?
The term "arbitrary" is to be understood as defining a family of
stress distributions that are preselected but still conform to equilibrium
requirements and boundary conditions. To answer this question we start by
making the following two basic assumptions:
(1) the classical equations of linear elasticity are valid in this
application, and
(2) the mechanical properties of the continuum can be expressed as
s atial functions.
Following from these assumptions the well known governing relations of gener-
alized plane stress, given in rectangular coordinates are as follows:
Equilibrium Equations
äa äcr xy ox T by + X
da Ba ~L + r-2 + Y »= 0 ,
öx öy '
Strain-Displacement Equations
(v/
II
du öv x dx ' y dy
dv . du xy dx dy *
(2)
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Compatibility Equation
^2 .2 .2 de de de
^ 2 ^ . 2 dxdy dy dx
(3)
■^ *■ • ■ L"
^
Assuming that the continuum exhibits orthotronic material proper-
ties and neglecting time and strain rate effects, Lhe constitutive equations
can be expressed as generalized Hooke's Law as
e X 'all a12 0 ' 'V V
e y
■ a21 a22 0 ay
+ < a;T
c 0 0 a66 4 .^
. o J (4)
f]
where a^ and a?2 are the coefficients of thermal expansion and T represents
the temperature difference distribution. The relations defined by Equations
(4) are valid when the axes of the material properties of the continuum are
coincident with the axes selected for the differential equations of the
problem. If the axes are not coincident, except for the z-axes, and the
other two axes of the material properties are rotated about the z-axis through
some angle, «(, in relation to the geometric axes, then more complicated
relations between the stresses, temperature and strains are developed.
Lekhnitskii [12] pr.sents these relationships in some detail. For th^ gener-
alized plane stress case in point the material coefficients are related to the
two axis system by
I
1
[!
II v
- —
•u •'u "'21
=
''22
■v
4 2 2 2 2 4 .2 2" m mn mn n 4Lin
m2„2 J* • 11,11 m n m n m n -4m n
m2n2 J* 4 2 2.22 n> n n m m n -4m n
2 2 2 2 4 ,22 mn tn n m 4m n
22 22 22 22 ,22 m n -ran -ran ran (m -n )
'11
'12
'21
l22
'66
(5)
where ra • cos rf, and n ■ sin i.
No generality will be lost by continuing with the constitutive
equation as given by Equations (4). Substituting Equations (4) into (3),
«- ^v *m im "
assuming that the a,,'s are spatially dependent and carrying out the required
differentiation yields:
{a12
^2 .2 .2 ,2 .2 0 C" _5si 4
&g« , Lfs a ill 66 dxdy all . 2 + an .2 a22 . + a.
dy
+ ^ (flfjl) + ^-J (o2T)} öy
öy
äx2
öx öx
{■ ä2a 11
9 + öy 2W
X
2 2 2 9 a12 ^ a21 ö a22
ay dx öx
2 2 ö a66 v , ^ all ÖCJx . , a +2 —: • t r 2 ^12 . day , 2
da21 OCT.
öxöy xy öy öy öy öy dx öx
öa2? öa^ aa66 *,„ **u ^xvl n
öx öx öx öy öy öx J (6)
Assume that the body forces have a potential, V, such that
I!
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öx
öy '
and choose Airy's stress function, Y, in the form
ay
(7)
.^4.V
y ax2 (8)
2 ö Y
CTxy " öxöy
^^
^
Making the appropriate substitutions into Equation (6) yields
( ox äx dy ay
Ä . ._ . . d2v . Ö2
+ (all + al2> ft + <a22 + ^l) 2 + H ^1T) + S ^2T>} äy dx dy öx J
I ^o a,, d a K ,i a^ + b2a ^ ^
öy~ dx' öy' N dy2 dx2 ' dx2
9 a,, ^2.,. ^2
dx ÖXÖ
2 2 2
f ' ti; + 71 (a22 + «ll) ' V + S (all + a12) * V 2
+ S; (2a12 - a66) S3Y
öx öy + fc <2a21 •w>
)3Y öxay2 ay ^3
JX % ax My^u^^'-^^^a^^-i}-0 (9)
D 0 I i i! I \
and, of course, the equilibrium equations are met exactly.
Notice that in Equations (4), (5), and (9), a12 has not been
equated to a21. Norm' lly, under the limit of small displacement theory
dealing with linearly elastic materials which are conservative, the material
coefficient matrix, defined in Equation (4), wou'J be symmetric and a.- would
equal a^. However, some recent work by Bert and ^uess [13], among others,
shows that there exists experimentally derived data which indicate that for
some types of composite materials exhibiting orthotropic properties the
material coefficient matrix is not symmetric and a is not equal to a. .
To limit the growing complexity of this work, the material coefficient matrix
will be taken as symmetric, at least for the initial phase of this effort.
Equation (9) can also be expressed in polar coordinates as
follows:
(Note: Here, the a.., are in polar coordinates and are
not equal to the a.., for rectangular coordinates.) lj 8
—
■ »-P T ■^ ^■p
^
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(fa ä\ + 2 121 äi _ ^1 A . il 8! . au aS o« är r dr r "' r a«
i) . >> / . 2,.2 ar as'
,2.u ♦ 2ai2 4. ^
^ r^ -ae
2a
^]
+ fa ä-r + I 21 : allNi a- 4. fU a2 1 A, •! or ,2
r 09'
a,.. .2
1 dr Lr ar2 r2 arae2 " r3 ae2 ar "~^~J
09 Lr2 ar2ae r3 öe J ör L2 ^ + r dr2 + 2 äT + "f J
r Lr3ae2 r2*r *■> 09 I 4,fl3 + r3ä7ä? + -2ärJ
aa "a^
aa "a MIX dfi Ip a3Y 1 . ^66
r Lr3 ae2 ' r2 arae2j ^ 1 Ä i- &. * J^ a3Y 1
.r4 ae ■ r3 arae + j ZXH r ar ae-
ar2 Lr2ae2 r ör rJ ae2 U2 ar2 r ^ :r- V eft
2 <.2a 22 a2a
ar "-ar'
a2.
Lar2 9j ae2 '-r^fl2 sar^^J
arae [2 arae " 3 aeJ J " 0 (10)
———— ■
^: r*m ^mm
T?
with / being the stress function defined by
^r r 8r 2 ^rt2 + vr
r dB
'• ■3+T« (ID
CTre or Vr 09/ '
Equations (11) satisfy the equilibrLura equations formulated in polar coor-
dinates for plane stress which are
or r 99 r R = 0
!Zr9 + l^+2^i + 0=o
(12)
In order for Equations (11) to meet Equation (12) exactly the body functions
V and V must be defined as r 9
fl \ ' h (rVr) " R
99 U '
(13)
thus putting a rather narrow interpretation on the body forces.
In the classical approach to the problem of orthotropic plane
elasticity, the coefficients a., are either constants, as in the case of the
homogeneous condition, or as in some rare cases, are special functions of
position. In the first case all the terms within the second set of large
braces, 1 |, in Equations (9) and (10) are zero leaving the remainder of
these two equations in the form of the well-known, homogeneous, orthotropic.
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10
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i !
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compatibility equations in terms of the stress function Y. In the second
case all the partial derivatives of the a, 's which appear in these second
sets of braces are capable of being evaluated, resulting in extremely compli-
cated fourth order partial differential equations wirh variable coefficients.
Proceeding along classical lii.es, these equations must be solved for Y which
c^ncains arbitrary constants of integration. These constants are then deter-
mined by evaluating Y in terms of the stresses on the boundaries.
Suppose it is assumed that the material coefficients, the a./s. ij •
aie unknown but that the stress function Y is a fully defined function of the
spatial coordinates. That is, the stresses throughout the body as well as on
the boundary are known. In such a case. Equations (9) and (10) reduce to
second order partial differential equations with variable coefficients, in
Leras of the a 's. The solution of these equations and the resultant deter-
mination of the magnitude and distribution of the material properties throughout
the body is defined in this worV as design synthesis.
It is believed that Equations (9) and (10) have not been previously
published. Bert [14] derived an equation similar to (10) wherein he reduced
the unknown material property coefficients from four to one. His formulation
is as follows:
c r A .l^Y e_iVp_^I+ 2(c-v) A U,.4 r . 3 2 . 2 ^ 3 or 2 „ 2iQ2 or ar r ar r r or 89
. 2(c-v) B3Y + 2(c-v+e) ^Y + e_ A 1
r3 örö92 r4 M2 r4 Ö94 J
3 SJY + dSr ^1 2^Ö?I e dY 2(c-v)
dr L ,. 3 r .2 " 2 dr 2 ^2 or or r r ^rdb r
.2„ . : 2(c-vHe A 1 4, A P-^- £ dl &. äV]
r3 ö92j dr2 ^r2 " r ör " r2 d92J
LV .2 " r ör " 2 ZJ '^V or r oo
^S^^-^blK)-- 2 ^
09 ^ (14)
m^
■^r •^ i
ii
ji
ii
v,ith the thermal terms oiritted. Equation (14) Is equal to Equation (10)
with the following identities:
S = a
ii
I
!
i
(15)
22
e - a11/a22
c ^ a66/2a22
vH " a12/a22
with S being the dependent variable and e, c, and v being fixed ratios. It
is anticipated that most real composite materials will exhibit material
property characteristics as defined by Equations (15).
The solution of equations such as (9) and (10) where there are 4
or more independent a.,'s, even where these a..'s are assumed independent of
temperature, is quite difficult but certainly not impossible. The simpli-
fying assumption made in Equations (15) leading to the formulation of Equation
(14) reduces the problem to only one unknown parameter.
J ""«i imiM UTi
T: ^* ^m T 12
APPLICATIONS
Rotattonally Symmetric rroblems
Example 1: The pressurized annular disk, internal pressure: Consider a
pressurized annular disk as shown in Figure 1. In the case where the ring
is Isotropie and homogeneous, the stress distribution is as developed by
Lame (1852) and is given by (c.f., Ref. [I5l, p. 60),
a *- r (VO 1 p
k - 1 p
(16)
where p » r/b, and k ■ b/a. These relations lead to the following con-
clusions:
(1) lcJ>!cr | for all p and all k ö r
(2) (aQ) occurs at the inner boundary (p ^ 1/k), and thus
I (Vmax P(4^)>P \- - v (17)
From the stresses so generated, which for the homogeneous ring, are quite
Independent of the material properties, it is clear that the material is not
being used efficiently, particularly as the thickness ratio, k, increases.
For the homogeneous, orthotroplc annular disk Bienick et al. [16]
shows that the stresses are given by
-(f+1) , f-1 ar " cl c2r
79 * -(f+lK r f-1 -c.fr v +c0fr
(18)
1/2 where f ■ (aii/a22^ » an^ ci an^ c2 are determined from the boundary con-
ditions, in this case CT ■ -P at r ■ a and j ■ 0 at r - b. This work and
X ♦ \
\
J ^-^ ^^
^
13
I
II
I
ii
FIGURE i. PRESSURIZED ANNULAR DISK
-,„'
A^
_ ^T ■ v r_n—:—. . J *■ "N
I
ll! II
!
I Li '\.
T^-
14
vork done by Shaffer [17] show that as k gets large the limit of the stress
concentration approaches the orthotropy ratio, f, avid does so rapidly. Both
Bienlck and Shaffer deal with the nonhomogeneous case, but both assume a
special form of nonhomogeneity, mainly that
all " •ll ^
a„ - a90 r *22 *i22
a12 " a21 ' a12 r
where \ is ^al and a.., a22, and a^2 are constants. Both carry out optimi-
zation by varying \ and observing the results, but no attempt was made ^o
carry out design synthesis directly.
Consider now that the disk is composed of a heterogeneous, ortho-
tropic material where the material properties are undefined but are functions
of the radius of the disk. The equilibrium equation for this case reduces to
dT + ^T"1-0- (W)
The compatibility equation reduces to
lr ("e5 " er " 0 • (20)
and ehe stress strain relation for an orthotropic medium is
er " ^r + aI2CTe
ee " a12CTr + a22CTe
(21)
where a^^, aj«, and a22 are functions of r. Assuming a stress function
such that
■
■ '
^r -^ ^^^^i ^
^
15
'.-i' (22)
and substituting Equations (21) and (22) into Equation (20) and carrying out
the required differentiation results in
•22f,, + C,
22*i*22>f,4'<J*,l2-i2*ll>,r "0. r (23)
where the prime marks designate absolute derivatives with respect to r.
Equation (23) is a single differential equation with three independent
material coefficients as the operational parameters. When dealing with specific
materials they may be found to be completely independent or show some type of
defined relationship. As a first step we will assume that there does exist a
simple ratio relationship between them that is expressable as
a99 * a99 » a 99 B *' '22 22 22 22
a11 «^22, a'11 ■ V'lJ (24)
a12 " k2a22' a,12 ' k2al22
Making the appropriate substitutions in Equation (23) yields
fl [V +7^ Y] a'22 + [Y" +7Y, - -|Y] a22 - 0 . (25)
f
Thus if Y is known a22 is fully defined. Let us suppose that for effective
material utilization it is desired that a be constant throughout the
disk; i.e., a. " A0. Integrating the second of Equations (22) gives
Y - A r + B . o o (26)
Applying the boundary conditions that a (a) « - P and a (b) = 0 yields
J ^4 ■ It ^
^r 9m V 16
Pb S^T tp - n
P rl ,, (27)
k - 1
with p ■ r/b and k ■ b/a as before. It is interesting to compare Equations
(27) with Equations (16) and (17). From Equations (27) it can be seen that
if k is greater than 2, a becomes less than the pressure P, and a , which
equals P at p ■= 1/k, becomes the maximum normal stress in absolute value.
Thus for such a stress function and geometry there is no effective stress
concentration. This is, of course, never true for the homogeneous disk.
With the stress function now fully defined. Equation (25) becomes
[(l+k2) - ~|-] ,22 + [—r^Hr]* 22 (28)
which has the solution
II
0 r (i
ii ii
l22 •V»>MI/*a-v] (29)
where p ■ r/b
I - - kj/k,
k -k S L(l-k2)k2J •
and for the modified orthotropic condition as defined by Equations (21) and
(24)
a22 B a22 " EQ 9
aii h k, ■ c „ 1 a22 Er
(3o)
k .J» 2 "^ " ver ' k .
■ — ——
* -» ~^ rta^
^r^ . . jm ^&~*m Tf—
17
i i:
The inverse of Equation (29) or Ee(r) is shown plotted in Figure 2
for various k's while k- was held constant at 0.5. This figure shows the
strong dependency of the modulus distribution upon the orthotropy ratio k..
The ratio k» has a lesser effect as shown by Figure 3. Here the isotropic
case (k. ■ 1) is shown plotted for five values of k2.
Example 2. Pressurized annular disk, external pressure: Consider the
pressurized annular disk as shown in Figure 1 but with the pressure acting on
the OD rather than the ID as shown. If the stress criterion is retained; i.e.,
aft ■ constant, then the stress function becomes
k ■j [1 - pk]
9 . -a- [i . k] r k-1 p
(31)
I '• k-l
I
where q ■ external pressure
k •= b/a
p ■ r/b .
The compatibility equation (25) becomes
k„a (1-kJ k.a. [(i+k2) .-|-]a.22 + [-rL + -±-] a22 =0 (32)
Equation (32) is the same as Equation (28) except that a replaces b. The
solution of Equation (32) is
II l22 -V')M[ä*<l-vJ< (33)
where 0, §, and p are defined in Equation (29). Comparing Equations (33)
and (29) we see that they are not the same, differing by the quantity 1/k
■ ■ Mt riMM
-X — •^m ^^
18
i:
i'
!
!
I
1 !
i tl
1 .0
0.9
0.P
0. 7
0.6
tu V
0.««
0.3
0.?
0. 1
K, =Variable, E = Eg
K2=0.5
o. o l
o.o
K, = 0.01
o. ? o.«« o. f.
flÜOULUS V.S. BA01US
0. 8 1 .0
FIGURE 2. MODULUS VARIATION FOR ORTHOTROPIC DISK WITH PRESSURIZED I.D., WITH a- - Constant
0
■ ■ -^ MM
"^7 *9 w mm 7 19
II
r
r
I,
■
f (1 I
E
/ E n A
x
I .0
0. 1
o. e
0.7-
0 . 6
0 .5
0 .•*
0 . 3
0 . 2
0 . I
K, = 1.0 (Isotropie Cose) E = EQ= E r
Kg5 ü= Variable
o.o
K2=0.5
K2=0.4
K2=0.3
K2=0.2
K2 = 0.l
o. o o.? o.H 0.6
MODULUS V.S. RADIUS
O.P I . 0
FIGURE 3. MODULUS VARIATION FOR ISOTROPIC DISK WITH PRESSURIZED I.D., WITH ae - Constant
"^ - - i i ■>*
■^7 m T?
X
20
!
within the second factor. Thus the distribution of the material parameters
through the thickness of the disk must be different as compared to the
internally pressurized disk. This is shown in Figures 4 and 5. Also notice
that for the internally pressurized disk the hoop stress aQ tends to zero as
k tends to infinity while the limit on CTQ for the externally pressurized disk
is q, the pressur?. Thus for the externally pressurized case the material is
not being as effectively utilized and perhaps some other criterion might more suitably apply.
If the annulus has both Internal and external pressure and the same
stress criterion is applied, i.e., C7Q « constant, then
^ ^ kTT [p " qk] »
(34)
and the material property variation is given by
*•*.<*"&* aM&l (35)
I I I 0 II
Example 3: The rotating disk: Consider the rotating, uniform thickness
annular disk as shown in Figure 6. The equilibrium equation governing this case is
fc (rt7r) " CT9 + g ^ - 0 ' (36)
The compatibility equation in terms of strain and the stress-strain relations
for an orthotropic material are given by Equations (20) and (21), respectively.
Substituting Equation (21) into (20) results in
0-a,12CTr + a12CT'r + a,22<7e + a22CT,e + r [(a12-all>CTr + <a22-a12)(V' (37)
i^te
^N Pi l^-ü w w~- *.——
21
!
r
n
i i
i
o
ii
/ E n A
X
1 .0
0.9
o.e
0.7-
0 . 6
0.5-
0.««
0.3-
0.2 -
0. I -
0.0
0.0 0.1 0.6
nODULUS V.S. RADIUS
I . 0
FIGURE 4. MODULUS VARIATION FOR ORTHOTROPIC DISK WITH PRESSURIZED O.D., WITH 0o ■ Constant
«U
-
"T ^m ■ m •Pi 7 22
r
i
I
I
E
/ E n A X
I . 0
0.9-
o. e -
o . r
o . 6
0.5-
o. H
0 . 3
0 . 2
0. I -
0 . 0 0 . 0 0.2 OH 0.6
nnouius vs. RADIUS
0 . * I . 0
FIGURE 5. MODULUS VARIATION FOR ISOTROPIC DISK WITH PRESSURIZED 0. D., WITH aa - CONSTANT
n
"^7 —-^
0
in 1
/ n
i
i
23
■J-Jl ZÄ l' -
Boundary Stress
FIGURE 6. ROTATING ANNULAR DISK
11 J
)
\
^.^^L. ^^
II M ^7 . . 9m p—
^
I 24
Assume the stress function
i.
f + — CD r
(38)
g
i
It is possible that the material density, y, could be a function of tie
radius. At this stage such a condition introduces an unnecessary complica-
tion and will not be considered. This being the case the governing
compatibility equation becomes
r Y 2 2 / , , a22 a12\ n (39)
Again, for simplicity, the modified orthotropy relations as defined by
Equations (24) will be adapted making Equation (39) take the form
0 - a'22[r+^ Y+^ü2r2] + a22[r'+f-"f+ ^a)2r(3-k2)] (40)
I
The stress criterion will be the same as before, i.e., aö B
constant. Using this criterion and operating on the second of Equations (38)
together with the assumed boundary conditions that a (a) = a (b) ■ 0 yields
Yv2 / k \ r, 1 11 2 ^ ^ -r-3rfc)L1-^-^ + ^3-p +kJ (41)
II fi
Yv '9 "It (Ä) M
.J l l «I*
^ —■-^
25
I i
0 I
where p • r/b
k » b/a
v B ujb * tip velocity
u) » rotational velocity, radians
g •« acceleration of gravity.
Here we note that in the limits, when k -» 1
w2
which is the stress in a rotating thin ring, and when k
very small)
Vv2
(42)
-> ■ (i.e., when a
(43)
which is smaller than exists for the Isotropie, homogeneous case by the
ratio
i 1
i [ II 0 II
u
where aQ • hoop stress for heterogeneous case öl
a« ■ hoop stress for Isotropie, homogeneous case o
v » Poisson's ratio for Isotropie, homogeneous ease .
Substituting Equation (41) into Equation (40) and carrying out the
required differentiation yields
where
, fAr 4- Br 4- D ~| 22 " I 4 , „ 2 . „ J
Fr + Gt + Hr- a22 -0 (45)
(kj - 3k2)
B - (1 - k.) rbi^ lJ L b
3-,
^^
^ .— > « ^
26
D - k^abHb+a)
F - k K3 3
H - k2(ab)(b+a)
!
i
1° 1
1
II
Equation (45) is not readily solvable in closed form. Before proceeding
to the solution of Equation (45) by some numerical means, it is convenient
to digress at this point to discuss another stress criterion which may be
applied to the implementation of design synthesis as defined in this
work. This criterion is that through the body of the disk, the in-plane
shear stress, T, is to be a constant. For a body of revolution, acted
upon by symmetric loads,
or
t - (ae-ar)/2
T - oQ-o,. - Constant ■ C or o (46)
Applying this concition to Equations (38) and integrating leads to
1 2 2
^ - C £nr - ^-^- + Cn r o 2g 1 , 2 2
4" - C £nr + C - 4 Yu r + C, o o 2 g 1
r R MU »
requiring that on the boundaries;
a (b) - o r ' o
a (a) - o, rv i
leads to the following relations.
(47)
A_
•^m
^
27
i f i 1
!'
i
I
i
I
0 I 1
If
i 2 2
a - ±y = C Inv - Y^ r r r o 2g + C,
2 2 yi + y"> r
g
2 2 Co<tnr + l) -■i|i^+C1
£n {<<'o-0l' * »Sift2-2)}
£n - a
!o.J!,nb - a Zna I i o
2
2g (b2Jina - a2£nb) (48)
These relations, as well as those defined by Equation (41) (a. = constant)
were applied to Equation (40) and were solved numerically. A digital
computer program, Determination Of Modulus Variation 1 (D0M0V1), was
structured for the solution of these sets of equations. A finite dif-
ference method of solution as developed by Manson[l8] was used for the
calculation algorithm. This program, which is detailed in the Appendices
was used to solve several problems as follows.
• The modulus variation for an orthotropic disk with
a pressurized I.D. for o - constant (disk not rotating)
• The modulus variation for an Isotropie disk with
pressurized I.D. for.o - constant (disk not rotating)
• The modulus variation for an Isotropie disk with
pressurized O.D. for o - constant (disk not rotating)
• The modulus variation for a rotating orthotropic disk
with no edge loads for OL ■ constant o
• The modulus variation ior a rotating orthotropic disk
(Poisson's ratio variable) with no edge loads for
a = constant I
• The modulus variation for a rotating orthotropic disk
with no edge load for constant in-plane shear stress
• The modulus variation for a rotating orthotropic disk
(Poisson's ratio variable) with no edge loads for
constant in-plane shear stress
.-
V
A.
» ■ ■■ p
^
I 28
The modulus variation for a rotating orthotropic
disc with 60.0 inch O.D., 6.0-inch I.D., turning
at 2,000 RPM, stressed on the O.D. with an
uniform edge load of 10,000 psi for o = constant
The modulus variation for a rotating orthotropic
disk with 60.0-inch O.D. 6.0-inch I.D., turning
at 4,000 RPM, stressed on the O.D. with a uniform
edge load of 12,000 psi for o«" constant.
r
i:
f
The solutions for the first four problems were checked by use of the closed
form solutions. Equations (29) and (33) and are those shown plotted in
Figure 2, 3, A, and 5. The solution to the remaining problems are shown
plotted in Figures 7 through 12. It should be noted here that when
boundary tractions as wel_ as body loads are applied to a rotating disk,
the solution is specific as regards the magnitude of the edge loads and
the rotational velocity of the disk. HowevPt, where only one type of load
is imposed, then the solution is generalized and is independent of the
magnitude of the load. (Note: The stresses remain directly dependent
on the load.) This solution dependence upon load is shown very clearly
by comparing Figures 11 and 12. Here the modulus variation is shown to
change with the change of the boundary tractions and the rotational
velocity. All these rigures have been non-dimensionalized by the expediency
of plotting E/E and R/B where r 0 max
E Ea, the modulus of elasticity in the 9 direction. 6
E ■ the maximum E„ calculated in the body of the disk, max 6
R ■ the radius of the point in the disk at which the
modulus is being calculated.
B ■ the outer radius of the disk.
The computer output for the curves plotted in Figure 12 is also found here
in Appendix C. For this case, o. equals 28,45A psi throughout the disk.
For an Isotropie, homogenious, flat-angular disk of the same material
■
^cr
ii
E
/ E «
I 0
O.*»
0.8-
0. T
0.6-
0 .5
O.t
0. 3
0. 2
0. I -
0.0 0. 0 0.2 0.I 0.6 0.^
• /• nnouL us v.5. RAOI us
1 . o
I \
FIGURE 7. MODULUS VARIATION FOR ROTATING ORTHOTROPIC DISK WITH EDGE LOADS, WITH o. - CONSTANT
B
NO
—————— ,
• ■ Mlii
30
II II r r r i
fi ii i fi J
E
/ E n A
x
I .0
0. V
0 . 8
0. 7
0.6
0 .5
0.««
0 . 3
0. 2
0 . 1
0.0 0. 0
K2=0.4
K2=0.5
K, = 2.5.E = EÖ
K2 = ü = Variable
-L.
0.2 0 . H 0.6
nnnuLus 9.%. RAOIus
o e i . o
FIGURE 8. MODULUS VARIATION FOR ROTATING ORTHOTROPIC DISK, ORTHOTROPIC RATIO OF 2.5 WITH NO EDGE LOADS AND oe - CONSTANT
•
\. 1
11
Mb
1
1
i
0 i I J
I . 0
0 . «»
0 . 8
0.7-
0 . 6
0 .5
0 .t
0. 3
0 . ?
0 . I
0 .0 0.0 0 . ? 0 . M 0.6
R/B nonuLUS vs. BADIUS
o . e i . o
FIGURE 9. MODULUS VARIATION FOR ROTATING ORTHOTROPIC DISK WITH NO EDGE LOADS, AND WITH CONSTANT IN-PLANE
SHEAR STRESS
I
I
D i 0
i
E
/ E n A X
I . 0
0 .9
0.8
0 . 7
0 . 6
0 .5
O.t -
0 . 3
0. 2
0 . I
0.0 0.0
-L -L
K, =2.5 E=Ee
K2-u= Variable
i
0.2 0.1 0.6 R/B
«OOULUS V.S. RADIUS
0 . P I . 0
FIGURE 10. MODULUS VARIATION FOR ROTATING ORTHOTROPIC DISK ORTHOTROPIC RATIO OF 2.5 WITH NO EDGE LOADS AND WITH CONSTANT IN-PLANE SHEAR STRESS
\
■ ■ -^
I .0
I
I
I
I
0 . 2 0.1 0.6
nnouLus v.s. RADIUS
0 . 8 » .0
FIGURE 11. MODULUS VARIATION FOR ORTHOTROPIC ROTATING»DISK STRESSED ON O.D. RPM - 2000; DENSITY «0.1 LB/IN. ; STRESS ON O.D. - 10,000 PSI: O.D. - 60.0 INCHES: I.D. = 6.0 INCHES, 0o - CONSTANT o
^JT
I I !
1 1 I
i
!
I
E
I E
1 .0
0.9
0.8
0.7-
0.6
0.5
O.t
0. 3
0.2
0. I
0.0 0.0 0.2 0.1 0.6
MODULUS V.S. RADIUS
0 . 8 1 .0
FIGURE 12. MOPULUS VARIATION FOR ORTHOTROPIC ROTATING DISK STRESSED ON O.D.: RPM - A000; DENSITY = 0.1 LB/IN.J; STRESS ON O.D. - 12,000 PSI; 0. D. - 60.0 INCHES; I.D. - 6.0 INCHES, oa - CONSTANT 8
^
35
I
I
!
1
density ofl would equal
b -a
ae - 35,810 + 24,2A2 = 60,052 psi.
Thus, by the use of design synthesis, the maximum stress in such a disk
can be reduced by a factor of 2.11. Further, Figure 12 shows that this
can be achieved with a modulus variation through the disk of approximately
2.5 to 1 for a material that exhibits and orthotropic ratio k, of 2.5
(the similarity of the two ratios is coincidental and bears no significance).
Non-Svtnnetric Problems
Example 4: Small hole in an infinite plate. Figure 13 represents a small
hole in an infinite plate which is subjected to a uniform tensile stress, P,
in the x-direction. For the homogeneous, Isotropie condition, the stress
distribution around the hole is well known as given by Timoshenko [15] as
r r r
2 A. ae -li1*3^) -K1^) cos2e
r r (50)
P / 3a 2a \ ar9 ' " 2 I1 " "T + TV 8in 2e '
The maximum stress occurs at r » a, 9 • (TT/2,3TT/2) , and is
max 9 r-0,e™n/2
For the homogeneous, anisotropic >. dition work by Green and Zerna [19],
Hearmon [20], Savin [21], LeckhnisKii [12], and (as directly applied to
composites) by Gveszczuk [22], shows that the maximum stress at the hole is
usually greater than for the Isotropie case and can reach values as high as 9P,
■ 1 Jiw
I
I i i
I
i
I
0 !
FIGURE 13. SMALL HOLE IN INFINITE PLATE SUBJECTED TO UNIFORM, UNIAXIAL TENSILE STRESS FIELD
■■»■■»■■■——————
I 37
Referring to Figure 13, consider the portion of the plate within
a concentric circle of radius b, large in comparison with a. It can safely
be assumed that the stresses at radius b are effectively the same as in a
plate without the hole and can be given by
I
(a) «T P (1 + cos 2 9) r r«b *
(a.) « - 7 P sin 2 9 . rö r«b z
(51)
From Equation (50) it can be seen that
((7Ö) WT P (1 - cos 2 9).
It seems reasonable, then, to choose a stress criterion for the plate
rQ ■ function of 9 = j P (1 - cos 2 9). (52)
From the second of Equations (11), with V equal to zero the stress function
becomes
Y -^ Pr2 (1 - cos 2 9) * f1(9)r + f2(9) . (53)
where f1(9) and f2(9) are functions of 9 only. Applying the first and third
of Equations (11) to ¥ results in
a_ am (i dr Vr ae/ "-CT_ft-TPsin29--ir 1 df2(9)
'r9 2
1
2 d9
1 d2f1(9) . d2f2(9) .
(5A)
rt^ '^H^r «lP(l + cos29) ^^^-^-^-—^-^ t^ r do ard9 r d9
Applying the boundary conditions
I (•J - (arö) - 0 » r-=a rö r«a
yields
38
I i
f2(e) » - ^- cos 2 9
d fl(e) P hr— + f^O) - - f ll + 3 cos 29] .
de2 1
(55)
Choosing as a particular solution to the second of Equations (55)
)
i
yields
and results in
f1(9) " Cj^ + C2 cos 2 3
fl(e) - " f^ [1 ' co829l
Y - | {(r2 - 2ar) - (r - a )2 cos 2 9}
CTr -Kl-^l-cos 29+ (l-^)cos29]
P '9 2 ■ 1 ■ "ö . - - cos 2 9)
^--K1-^)81"29-
(56)
(57)
i
39
The stress function, y, as defined by the first of Equations (57) was
applied to Equation (10) with the body functions and temperature difference
taken as zero. In order to implement a solution, the following relation-
ships among the material coefficients i re assumed.
i
all " kla22
a22 " a22
I
I
a12 " k2a22
a66 ' k3a22 '
A finite-difference algorithm was structured to carry out the solution.
The method of solution applied was that usually referred to as the
"relaxation method" which is discussed in detail by Shaw[23], Hildebrand[25],
Allen[26], and Richtmyer[27].
Due to symmetry, only one-fourth of the plate was modeled. A
square mesh was used and the quarter plate separated into a square array
of 61 x 61 nodes. It was assumed that the circular hole has a radius of
1.5 Inches and the mesh distance, h, the distance between nodes, is 0.25
Inches. The width of the quarter plate model, thus became 15.0 inches.
This gave a b to a ratio (see Figure 13) of 10:1, thus minimizing the
far-field effects of the outer boundaries upon the stress around the hole.
This plate model is shown in Figure 14. Rectangular coordinates were
employed and the differential equation that was differenced was that shown
in Equation (9). Equations (57) were converted to rectangular coordinates
for input to this program. In the solution procedure, it is necessary
to assume the values of the modulus parameter a „ at all boundary nodes.
The solutions were then to proceed in an iterative manner until the
values for a.« were determined at each interior node in the model. No
success was attained by this method. The program never was able to
converge to a solution. In fact, a strongly divergent tendency was noted
(i.e., each iteration on the a^'s at each node point was markedly greater
^_JL_
AO
!
I i I i f
1
!
Outer Boundary Points
eg 531 6QI 58,1 Line of Symmetry
7,1*6,1 5.1 4.1 3,1 2,1
Inner Boundary Points
FIGURE 14. TWO-DIMENSIONAL FINITE DIFFERENCE MODEL
_J
41
on each iteration in the relaxation process than the iteration which it
succeeded, regardless of the boundary values assumed.) Extensive investi-
gations indicated that thera was no error in the solution process. It
appears that there must be some other governing relations which hav as
)■
(
I I I !
i
I
ji J
yet not been expressed.
As a result, no solutions to tne two-dimensional problem in
design synthesis have as yet been accomplished.
-
■ * -^ MM
42
Limits and Other Considerations
In the preceedlug discussion, it was pointed out that a solution-
for the two-dimensional problem could not be achieved. This raised ques-
tions as to whether all governing equations have been generated. In the
application of the one-dimensional program, D0M0V1, ir was found that
under certain conditions, the only mathematical solution developed for
the modulus variation in a disk would require that at some points in the
disk body, the value of the modulus must be negative. Though this is
perfectly acceptable from a mathematical viewpoint, it obviously cannot
be implemented from the viewpoint of real materials. To illustrate this,
consider the case of the rotating annular disk with boundary tractions
when subject to the stress criterion that the in-plane shear stresses
throughout the disk must be constant. From Equations (48)
2 2 a - C (Anr + 1) - ^^-^ + C. ti o o 2g 1
i
!
I
1.1 II
2 2 2 0 ■ e tar - Y o y + ci TO 2g 1
C - JL- {(o -a.) + *f-(b2-a2)} (48) 0 £n b I o i 2g
a
1 / 9 9^ C. - ^ ia.inh-a £na - **£- (b Jlna-a Ä,nb)?. 1 b \ i o 2g )
a
Take the condition where the disk is not rotating; i.e., u ■ 0, then o -o..
An — a a.Änb-o £na
C, - ^ r-* (59) 1 in^ a
o - C (Ä,nr+1) + C. 9o 1
o - C Anr + C. . (60) r o 1
• i M>i
^s TTH • ■ P
^
43
Substituting Equation (59) in (60) yields
T \(.a -a.)(inr+l) + a.£nb-a Una}
0 Ü f
\
or - r Uao-a )£nr + a J,nb-a £na> .
a
Now, further assume that a - 0, a - -P . Then, Equation (61) become
(61)
in — a
^ I" n- (62)
Note that for all values of r <^ b, a < -P in the algebraic sense.
However, if - > e, (2.71828), then at r - a. In - > 1 and aQ is negative. a a 6
Defining the displacement at r = a as
Va) -{m '"e-^r1) lr.
V.) -utr tt u •ta i*v *• «'I u (r-a) " E^iT^b [1 " ^a * (1-v)] (63)
Ä,n - a
From Equation (63), if In — > -j— then the term in the brackets is negative
and U or E(8) must be negative. If either is negative, then the system
must do negative work, which is not possible for real materials. This,
of course, can be overcome in a numerical sense by simply requiring that
only those solutions are acceptable which yield a material coefficient
matrix that is positive definite at each point in the body; i.e., the
following conditions must all be met:
M ' -L ^
-^-
A4
all^0
alla12 " a12 ^ 0
alla22a66 " a12a66 i 0
(64)
Thus, It is clear that the "arbitrariness" of any stress conditions selected
are restricted to more than conforming to the equilibrium equations and
the boundary conditions.
Two Dimensional Boundary Considerations
!
It is instructive to approach this problem from the viewpoint
of the calculus of variations. Consider the total strain energy in a
stretched plate uf unit thickness:
i;
i M
I
f
r fi
U - //. U dxdy , A O
where
U = strain density at a point
■■T-[oe +oe +T Y ]. 2 xx y y xy xy
Assume a Hookien material, neglecting time and temperature effects,
e ■ 3,-0 + a,-o x 11 x 12 y
e ■ a.-o + a-.o y 12 x 22 y
xy 66 xy
and a stress function
(65)
(66)
■ J -^
_ ' _^^ ^
45
2 o. ay2
3x2 (67)
[ r r
i
xy ax3y
Such Chat Equation (65) becomes
U - // i[-u($)2-u($X$) + a 22 (S) + a66 (|S7) J^^ .)"}
Following the method of Weins took [ 28 ], the extremization of (68) is
affected by forming the integral 1(6) by replacing f in the integral
of (68) by
4 - nx.y) + 6n(x,y) ,
(68)
(69)
Ma
I !
r !
I I I I
'■■•.
1 X J
\
where ^(x.y) is the actual extremizlng function and n(x,y) is an ar-
bitrary function that is twice continuously differentiable. Then, the
integral 1(6) is an extremum for 6 - 0, so that
I'CO) - 0 .
Writing f(y , y , y ) 0 xx yy' xy the integrand of Equation (68), and here the
subscripts refer to differentiation with respect to x and y. Then,
according to (68) and using (69) to compute
ay XX
3t
36 - n J2. 3f
xx* 36 JÜL
'yy* 36 xy
* I -^
— III!
M^M^Mi
(70)
(71)
i HUM
j m
Tl
A6
and I'(6) Is formed and 6 Is set to zero, resulting In
I,(0> " "A (#" \x + #" V + ^ V)dxdy = 0' (72) v xx yy xy v
according to Equation (70). Integrating by parts and employing Green's
theorem results In the transformation of Equation (72) to
//, \{im*im^m\ Mnt*kv)+^ (%)]"" 3f 1
- T n
dxdy
3f / y 3* 2 "x a* (
yy yy)
dx
, r >f 1 9f c) 'x a* 2 'y ay
{ xx ' xy
[» /»f \ . 1 3 / 9f \"| / .
r Noting that
0 . (73)
3f 324' ■'324'
xx 12(;y2)+a22l,3x2J
3f 34»
yy
3f 3*
xy
'11
l66
/32Y\. /As l3y
.2
^3x3y ' "Yxy
(74)
(Note: Confusing subscripts. Subscripts on strains do not refer to
differentiation.)
thus, the area Integral becomes
//. K 32e 32e
ax 3y
32Y
3x3yjj dxdy . (75)
■ ' Mfci ■*
^s ^
47
Since (73) must hold for arbitrary n, it must In particulrr hold for those
n for which n = nx = n «= 0 on the boundary C. For such n Equation (73)
reduces to the well-known compatibility equation as previously derived
(see Equation (3)) and the resulting (Equation (9)). For arbitrary n, n , n , x y
other than zero, the boundary relations must be derived from the remaining
line integral portion of Equation (73). The first part of the line inte-
gral becomes, along C on y - constant,
4 3all »2¥ 3^ 8y 2y2 ^V
+ Üü A + (a + I . j JL1. 3y . 2 + u12 + 2 a66; ! 2,
3x 3x 3y
iÜ66 2 3x )x3yj
I
r 32*. aM if sM 663x3yJ
The last two terms in Equ^ tion (76) are
- n. [Cx]-[\ky] • If these strains are not prescribed as zero, then along this portion of
the boundary n
fied is
n ■ 0, and the boundary equation which must be satis-
3a 11 3*4' a.^3* . 3a12 S2* ^ , 1 v 3 f
(76)
(77)
1 ^66 _£l 2 dx 3x3y 0. (78)
and a similar set of relations can be written for the second part of the
line integral. These equations have, at best, a very limited application
other than showing that other constraints do exist on the boundary.
Their limited application is due in great part to the choice of a rectan-
gular coordinate system. All attempts to date to express these relations
in terms of other coordinate systems, in particular those using generalized
normal and tangential components have been unsuccessful. More work must
be done along these lines.
111
■ ■ -*-"
•^m ^T
i r r !
I ■
I I 1
A8
Material Considerations
Finally, just a brief note on the possibility of achieving a
material with tallorable properties. It is sufficient here to say that
work by Adams and Tsai[29], Dimmock and Abrahams[30], Hewitt and Malherbe[3l]
Halpin[32], Halpin and Pagano[33], Kohn and Krumhansl[34j Tabaddor[35],
Fotlnich[36], Wang[37], and Fokin and Shermergor[38], among others, have
clearly established that material properties for composites of various
types can be established by a knowledge of the known properties of the
constituent materials, their orientation and their relative density.
J ■I ■■ WP w^-n-*
^s ——. jm
^
I I
I i I
I
I II
111
49
SUMMARY
Mathematical design synthesis has been shown to be possible In
certain specific applications. The selection of a design criterion In
the cases discussed, two dealing with stress distributions, and the
development of the material property distribution within a plane body
such that compatibility Is satisfied, appears to be a rational basis
of design for composite materials. Difficulty has been encountered In
the solution of two-dimensional problems employing this concept due to
the, as yet, undefined boundary restraint requirements which affects
the selection of the stress criterion to be met.
■ ■ -^
TT^ ^~
r I i -
I
I
I
I
1 fl
N
3u
REFERENCES
[I] Sayers, K. H., and Hanley, D. P., "Test and Analysis of Graphite- Fiber, Epoxy-Resin Composite", AlAA Paper No. 68-1036 (1968).
[2] Chen, C. H., and Cheng, S., "Mechanical Properties of Fiber Reinforced Composites", J. of Composite Materials, Vol. 1, p. 30 (1967).
[3] Hill, R. C, "Evaluation of Elastic Modulii of Bilaminate Filament Wound Composites", Experimental Mechanics, Vol. 8, No. 2, pp. 75-81 (Feb., 1968).
[4] Gaonkar, G. H., "Average Elastic Properties and Stress for Composite Continua". Collection of Technical Papers on Structures and Materials, AIAA/ASME 10th Structures, Structural Dynamics and Materials Conference, New Orleans. La., pp. 172-182 (Apr.il 14-16, 1969).
[5] Dong, S. B., et al., "Analysis of Structural Laminates", Air Force Report ARL-76 (Sept., 1961).
[6] Tsai, S. W., "Strength Characteristics of Composite Materials", NASA CR-224 (1965).
[7] Tsai, S. W., "Mechanics of Composite Materials^ Part I - Introduction, Part II - Theoretical Aspects", Air Force Materials Laboratory, AFML- TR-66-149, Part I (June, 1966), Part II (Nov., 1966).
[8] Tsai, S. W. and Azzi, V. D., "Strength of Laminated Composite Materials", AIAA Journal, Vol. 4, pp. 296-301 (Feb., 1966).
[9] Whitney, J. M., "The Effect of Transverse Shear Deformation on the Bending of Laminated Plates", J. of Composite*Materials, Vol. 3, pp. 534- 547 (July, 1969). *
[10] Whitney, J. M. and Leissa, A. W., "Analysis of Heterogeneous Anisotropie Plates", J. Applied Mechanics, Vol. 36, No. 2, pp. 261-266 (June, 1969).
[II] Whitney, J. M. and Leissa, A. W., "Analysis of Simply Supported Laminated Anisotropie Rectangular Plate", AIAA Journal, Vol. 8, No. 1, pp. 28-33 (Jan., 1970).
[12] Lekhnitskii, S. G., "Theory of Elasticity of an Anisotropie Elastic Body", Holden-Day, Inc., San Francisco, Calif., (1963).
[13] Bert, C. W. and Guess, T. R., "Mechanical Behavior of Carbon/Carbon Filamentary Composites", Composite Materials; Testing and Design (Second Conference). ASTM STP 497, American Society for Testing and Materials, pp. 89-106 (1972).
»
■ ■ -^ - — _^^_^^^——^^M^_^. _J
^T mmam W ^ -r—7-Tz _ •
11 [ [
!
Ifl
i 1 i !
I
51
[ 14] Bert, C. W., "Analysis of Nonhomogeneous Polar-Orthotropic Circular Plates of Varying Thickness", Bulletin No. 190, Engineering Experi- Ment Station, College of Engineering, The Ohio State University, Columbus, Ohio, Vol. 31, No. 2 (March, 1962).
[Li] Timoshenko, S. P. and Goodier, J. N., Theory of Elasticity, Second Edition, McGraw-Hill, New York (1951).
[16] Bieniek, M., et al. y "Nonhomogeneous Thick-Walled Cylinder Under Internal Pressure", ARS Journal, Vol. 32, pp. 1249-1255 (August, 1962),
[17] Shaffer, B. W., "Orthotropic Annular Disks in Plane Stress", J. of Applied Mechanics, Vol. 34, pp. 1027-1029 (Dec, 1967).
[18] Manson, S. S. , "Determination of Elastic Stresses in Gas-Turbine Disks", NACA Technical Report No. 871, Flight Propulsion Research Laboratory, National Advisory Committee for Aeronautics, Cleveland, Ohio, February 27, 1974.
[19] Green, A. E. and Zerna, W. , Theoretical Elasticity, Oxford at the Claredon Press (1954).
[20] Hearmon, R.F.S., Applied Anisotropie Elasticity, Oxford University, Press (1961). ^
[21] Savin, G. N. , Stress Concentration Atound Holes, Pergamon Press, New York, N. Y. (1961).
[22] Greszczuk, L. B. , "Stress Concentrations and Failure Criteria for Orthotropic Plates with Circular Openings", Composite Materials: Testing and Design (Second Conference), ASTM ST? 497, American Society for Testing and Materials, pp. 363-381 (1972).
[23] Shaw. F. S., "An Introduction to Relax-ition Methods", Dover Publications Inc., New York, 1953.
[24] Hildebrand, F. B., "Finite Difference Equations and Simulation", Prentice Hall, New York, 1968.
[25] Southwell, R. V., "Relaxation Methods in Theoretical Physics", Oxford at the Claredon Press, 1946.
[26] Allen, D. H. de G. , "Relaxation Methods", McGraw Hill, New York, 1954,
[27] Richtmyer, "Difference Methods for Initial-VAlue Problems", Inter- science Publishers, Inc., 1957.
[28] Weinstock, R. , "Calculus of Variations, With Applications to Physics and Engineering", McGraw Hill Book Company, New York, 1952.
•
^
52
I
r i
I I I r
II
[29] Adams and Tsai, "The Influence of Random Filament Packing on the Transverse Stiffness of Unidirectional composites", the Rand Corporation, RM-5608-PR(1968).
[30] Dimmock, J., and Abrahams, M., "Prediction of Composite Properties from Fibre and Matrix Properties". Composite, December, 1969, pp 87-93.
[31] Hewitt, R. L., and Malherbe, M. C., "An Approximation for the Longitudinal Shear Modulus of Continuous Fibre Composites", J. of Composite Materials, Vol 4 (April, 1970) pp 280-283.
[32] Halpin, J. C., "Stiffness and Expansion Estimates for Oriented Short Fiber Composites", J. of Composite Materials Vol. 3 (October, 1969) pp 732-734.
[33] Halpin, J. C, and Pagano, N. J. , "The L minate Approximation for Randomly Oriented Fibrous Composites", J of Composite Materials, Vol. 3 (October, 1969) pp 720-724.
[34] Kohn, W. , Krumhansl, J. S., Lee, E. H., "Varnational Methods for Dispersion Relations and Elastic Propertiets of Composite Materials", J. of Applied Mechanics, June, 1972, pp 327-336.
[35] Tabaddor, F., "A Note on Effective Moduli of Composite Materials", J. of Applied Mechanics, June, 1972, pp 598-560.
[36] Fotinich, 0. V., "Determining the Equivalent Modulus of Corded Rubber Test Pieces", Mechanics of Solids (Mekhanika Tverdogo Tela) Vol. 4, No. 3, 1969, pp 162-166.
[37] Wang. F. F. Y., "Calculation of the Elastic MoHnl1 of Binary Composites", Materials Scienre and Engineering, 7 (1971) pp 109-116.
[38] Fokin, A. G., and Shermergor, T. D. , "Calculation of Effective Elastic Moduli of Composite Materials with Multiphase Interactions Taken Into Consideration", Journal of Applied Mechanics and Technical Physics (Zhurmal Prikladnoi Mekhaniki i Tekhnicheskoi Fiziki) Vol. 10, No. 1 (1964), pp 51-57.
—
^^N . _j w
^
i I I
!
f 0 I I I I !
!
I
APPENDIX A-|
INPUT DATA FOR PROGRAM DOMOV1
Five cards constitute the input data for D0M0V1 for each
variation to be studied. These five cards represent three (3) data
sets. As many runs may be stacked behind the initial set as required,
The computer run will terminate when an End of File card is read.
Data Set 1 (Title Cards)
READ (5, 10) (ITIL(I, 1), I - 1, 24)
10 Format (8A10)
Any information can be placed on these three cards. It is
necessary to have three cards, although any can be left blank.
Data Set 2
READ (5, 20) RO, RI, RPM, DENS, N, IBOND, 1ST
20 Format (4F 15.0, 3115)
RO - Outer Radius, Inches
PI ■ Inner Radius, Inches
RPM ■ Revolution per minute of Disk
DENS - Material Density of Disk, lb/in.3
N ■ Finite Difference Increment and Printout Number
IBOND - Type of Boundary Condition on Inner Radius
0 * A stress condition
1 "■ A displacement condition
1ST - Type of Stress Criterion Chosen
1 - Constant theta stress
2 = Constant in-Plane shear stress
• i ■!>
'^^N ^m V
!
[
[ i
!
I
I I fl W I!
I
A-2
Data Set 3
If IBOND Equals 0,
READ (5, 40) SI, SO, ORTHO, PO, ETHETA
40 Format (2F 10.0, 3F 15.0)
SI - Radial Stress on Inner Radius
SO - Radial Stress on Outer Radius
ORTHO - Orthotropic Ratio, E /E0 r o PO - Poisson's Ratio
ETHETA - Modulus of Elasticity in Theta
Direction at the Inner Radius
If IBOND Equals 1
READ (5, 30) UI, SO, ORTHO, PO, ETHETA
30 Format (2F 10.0, 3F 15.0)
UI - Selected Radial Displacement of Inner Radius
i i i aJti ^fc_^_^_ M*
•^1 w Reproduced from best available copy.
APPENDIX B
LISTING OF PROGRAM DOMOVl
C
•')
PROGRftH 00V0V1 (INOUT.OUfUT.TAPFSsINPUT.TAPFftsOUTPUT)
OIMENSION RdMI.-TMruu-'KrOCCi:-)) ,/i22 (l D TJ ) , CC (ICC » , 00 (1C 2 I , t ... PHi(i::),PHio(i:a> ,PHiPP(ict:i,siGRiij:»,STGT(icci ,u(i:o)
_DIKZNSION 33(lJii) ,AA(1.21»££(1Ü3J c c
c c
COHION R,PHI,PHIP,OHIPP,SIGR,SIGT,U
_5-_PE.A0 (5,1: 10 FORMAT (BA IF(£Of,5)
15 RFAT (5,? 20 FORMAT iUF
IFdHOND-l -__Z5_PEAO (5,3.
3C FORMAT (2F C RO . C RI
X.. RPM . C OENS
.C __N C C C C C
.C C C C C C .C. c c c c c
_c c c c c c
IDOMO
1ST
UI SI SO ORTMO PO
.ETHE.T.A.
ROO EOC
PHI PHTP r>HIPP SIG«> SIGT U ETHET
)- (ITILU.l) UI 12.,15 1 RO,RI,RPH, 15.:,315) , ) 35,25,2^ ) UI,S0,r«9TH 1C.;,3F15..)
»0UTrP. PA = INflrR "»t sREV^LUTT sMATEojAL
-FINITE_C =TYPC Oc
0 = A ST 1 = A 01
sTYPI OF 1 = COMS
2. =. CONS sRAOTAL n »RAOTAL S sRAOTAL Z =0RTH0TO0 SPOIS«;OMS
. =.M00UL'JS. THt INHC
sRATIO OF SRATIP rm FOUND I*
=THE STRF . ..sTHE.FTRf;
»THp's^nr »RAOIAL t roipru^pf «THE PAOl «THE CALC
,l=UZkL
OENS,N.IOOMO,1ST
0,PO,ETHETA_
OIU';,INCHES OIUS, INCHES CMS 0Z0 MINUTE OF DISK
OENSITY OF niSK, LR./C IPFERCMCE. INCREMENT. NUM nnuNriARY CONDITION ON I RESS CONHITION . _ SpLACEMENT rONniTION ST^TSS CRITERION CHOSEN TANT THETA STRESS TANT IN-PLANE. SHEAR ST«» TSFLACEMENT OF TNNE« RA TPE3S ACTINC ON INNER P TOESS ACTTHF, ON OUTER 9 0IC RATIO, E-SU9-R/E-SU
PATIO IN THE THETA-R 0 Qr_-LASIIC T.l.Y_.IM_ J_H£„TM 0 RADIUS
TH". RADIUS OVER THE OU THE MODULUS OVER THE H TH" BOOY nfL THE_OISK
SS FUNCTION T P-RIVATTVr OF_THF_STR NO OERIVATIVE' CF'THE" ST TPESS .. R"MTIAL STRESS M 0ISoLACEMENT ULATEO riRCUMFFÄFNTIAL
mic INCH ?EF. »•'NER RADIUS
PRESENTLY
ESS HIUS AOTUS AOIUS 9-THETA IRECTION ETA DIR£CTIOM_AT.
TEP RADIUS AXIMUM MODULUS
rSS FUNCTION RESS FUNCTION
MODULUS AT FACH
OOM DOM DOM OOM OPH OOM. DOM DOM 00" DOM DOM DOM OOM OOM OOM 00^ OOM DOM OOM D0M
00" DOM DOM DOM. OOM DOM OOM 00M
DOM 00" PPM OP" OOM DOM OOM DO" 00" DO" OOM OOM DO" OOM OOM DO« OOM OP" OOM
1 ? j
(. 5
.. 6 7 3 o
10 11
.12 13 1« 15 16 17 .1! 1° 2'. II 22 2?
.2«. 25 26 27 29 2° 3: 31
33 3«. 30 3t 37 3? 39 <•:
«.2
^7
■ " -^
-^7 W
B-2
C CALCULftTnn POINT IN THE DISK OOH . _C .. A22 sl.C/FTHfT POM
GO TO us noM __35 f?FAn f5.£.Jl «JT^OfnoTHO^n^THFTi pnM
5: SI
UD FOo^ftT IZriQ.fitSriS,:) OOM 5?
OMEGAr(POM.OT,/,^..j» ppn 51.
G«?HO=<nENSI»(0MrGA»»2.:)/(3l6.i.) DOH 56 57
ETHET{l)=LTHrTA OOH RMlrPT OOM
5? 5°
OELi (f?0--?I)/(N-l» 00" 00 11, T = ?.»J pnM
5:
J=l-1 OOH PtTl=?<.)Wnri priM
62
T
5C CONTIVUE 00« 6!4
65
i: n
CALL ST-?-SS (N,iRA,r,^HrtSI,SO,UT,Ol,THO,PO,ETHETA,IPONO,IST,RI,RO» OOH _. DO 55 1=1.N . ........._. OOH
AA(I»=PMTP(I)MPn»PHT(in/(R(I))*{GRH0»(R(I)»»2.:n OOH -BB(I)=PHI0P(I)t?HIP<I»/R(I)-. (ORTHO^PHK 1))/(??(I)»*2.;i ♦(GRHO'Rd D0U
$)M3.:-O0)) 00" r,n(T) = mtT)/AAm nn-
66 67 6° 6 = 7: 71
55 CONTINUE OOH 00(1)=?.: nnn
72 7^
EEdlsC! OOH DO fir T=?.N noM
7<. 7t;
J=l-1 OOH 0D(I)=1. }*CC(l)»(n(I»-p|J))/(2.LI no-
76 77
EE(T» = I. '-cctj»•(--■(n-o«jn/2.: DOH . •Ctltl«fCCftl/QOIIIt«IAttfJII _._ .. . _ . DOM
ETHET(I)= l.i/A22(T» 00« ., _ feC CONTINU- nn"
7» 7°
i
EMAX=-,: noH 00 «i^ T=l fN pnH
1! 1'
EHAXrAMAXKETHTTd» ,EMAX) OOH 65 CONTINUE DOM
DO 7 C 1 = 1. N 00". ponm=Rm/Rn non
•»6
EocmsETMETm/pHftx noH U(T)=A??(T)»«>(T)»(SIGT r)*P0»STGPITH OOH
14
7C CONTINUE OOC. . P0=-1.?*D0 . DOH
MPITE (5,75» l|TKttttl«f«tUtM 00" . .. 75 FORMAT (1H1.8A1:/1H ,eAl.-'lH ,8Ai:///) _._. . .. . .__ ,. DO"
I«"(IPONO-1» i;»ti4«t*<l OOM 81 WRITE (5,95)P0,i'I.!:!:,H,5<)t^i,FTHrTA,.?O,OPTH0.0ENS.N . __ . DO"
91. 95
'—
- - 'I Jfc I
B-3
85 FORMAT (l;x,51Hr)ET^OHTV4TTO,l Of MODULUS Vft^TftTTON IN AN ANNULAR HIS SK.//,1X,1-HTNPUT 1ATA,/,iy,lCH»»»»»»»»»»,//,5X,15HOUTE9 RADIUS = , SFIC.I.,?/.ISHINNER "AOTUS = ,P1 J.««,2X,6HoPK s , Fl 0 , 2,//,5V , 3C HP AOI A
.-. »L STRESS, OUTFR PAOIUf. = , Fl t. ?. 2» ,3 CHRA PI AL STOfSS, INNE" RAOIUS t» ,Fi: .2,//,5X,25wrTH: TA AT INNER RAnjUS = , C0?:i5 . 7, 2X , 1 7w0nT SSONS
S RATIO = ,FlJ.9,2X,2:HQRTHOTPOPIC RATIO = ,F1C .«..//, 5X , iqH"ATE RI AL t DEHSITY = ,m.B,2X,25^NtJHPER OF RAOIAL "OINTS = ,13,///»
9C WRITE (6.95) - —•• - 95 FOPMAT (lX,ll"OUTt>UT PATA,/, 1X,11H»»»»»»»»»»»,//,7X,*)HRA0I"JS,PX,6U
SETHETA,'iXt''HSIr,1A-P,,<X,7H3IGMA-T,l',iX,«.HR/PO,lCX,«»HE/En,13X,7HU-SUn t_o,/,/x, ^M»»»»»», qy, fcH» • ••••,<»X, 7H»» •••••, 8X, 7^ ••••••»ICX.UH»»»»,
JllX««iM»»»».iliX.7-H»»».»l»S.»//J WPITf (6,1 JO) (R(I),rTMrTm , SI GR ( T ) , ST GT (I) , POT «I) ,E OC (I) , U (I) , 1 = 1
1C( FORMAT r>,,Fl..«»,2y,lPcl3.6,flX,F7.C,flX,F7.0,*;X,F«.5,6X,Pq.i«,6X,
tE12.6) GO TO IIS
.105 »"?ITE(6,liJ) PO,RI.Rn^,SO,UI,ETMETA,Pn,ORTHO,OENS,N 11C FORMAT (5Y,5^HO=:TrpMX'4ftTnN OF MOnuLUS VARIATION IN AN ANNULAR OIS
1:K,//,1X,1*HINDUT ^ATA./^X, ll:H»»»»»»*»»»,//,5X,15HOUTER R^OIUS = , SFlC.^X^SHIKNrTR PAHIIIS ■ , Fi 0 , (, , jy ,f MRP" = , Fl <; . 9,//, <; X, 3! HO AO IA
SL STRESS. OUTER RADIUS = , FIT.. 2 , 2X , 3?HRA?I AL DISPLACEMENT, INNE? R SAnilJS = ,Fi: .7,//,5X,?,;MTH-:TA AT INNER RADIUS = ,CPF15.7,2X, 17MDOI
tSSONS RATIO =_,F.i:.3 ,2X, ZIHDRTHOTOOPIC . RAT 10. =_, F13 . i.,//,^X, 1°HM AT tFPJAl DENSITY = ,Fi:.5,2X,26HNUMnER OF RAOIAL o0INTS = ,13,///)
.... GO TO 9: 115 CONTINUE
. .. GO TO 5 12P COWTINUE C^LL. EXLT -—
END
DOM 96 DOM 97 DOM qp
DO^ 99 DP« li? DOM 1 Jl OOM 102 OOM 133 DP** n«. OOM l!5 DOM i'.% DOM 1C7 DPM fin DOM 159 OOM 11C OOM 111 OOM 112 .DOM 113 POM liu DOM 115 DPy 116 OOM 117 POM 11.1 DOM 11° DOM i?: OOM 121 OOM 122 DOM 123 OOM 12<4 .DOM 125 DOM 12*
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SUD90UTINE STRE^SCN.Pft.GRHO.SI.SO.UI.ORTHCPO.ETHETA.IBONO.IST.RI, tRO» ...
oiHEKsioM R(in»,rTHET(ic:i,coc(i:o),a22(ico),cc(ioc),no(i2oi, $ PHH^-i» .DHI0(1-:» ,PHIPP(K:»,SIG9C10C ),SIGT11C;.) ,u<103) -i 1 ROB f 1C : 11 LT.XL (2 «111 DIMENSION RR(IJC),Äa«iji),EEIIP:»
_ c c COMMON R.0HI,PHIP,PMipp,siGR,SIGT,U
c c c c c r c _c. c c c c c -C c
THI5 SU^RfluTIN- CAIC'JL4TES THE STPfSS FUNCTTOK' AND ITS FTPS! ANO SECÖ"n DERIVAT IVESFOP THE FOLLOWING CONDITIONS, IF lOnHD = ] ANO IS! = 1 WE HAVE A STRESS CONniTION 0^ THE INNEP RADIUS AND A STRESS CRITERION OF CONSTANT TH^TA
STRESS JJL1PON0-« 9.ANO 1ST -_2_HL.HAVE A.ST'ESS CONDITION ON THEL
INNER RAPTU'; AMI 1 STRESS CPIT^inN OF CONSTANT IN-PLANr
. -. SHEAR STRESS . . IF IOONO = I ANO IST = 1 HE HAVE A OISPLACc-MENT CONDITION ON THE IHHf». PACIUS WITH_A STRESS CONOTION OF CONSTANT _
THETA ST^rSS. T-HESE..ARE . ItiE_rtlLY._C0bL1III9-NS_u30C-PA.MyEQ
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SECTION OeL A CONSTANT
THE .SUBROUTINE: CONTAINS THE. CALCULATIONS CIRCUMFERENTIAL STRESS CHNDTION
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*r. = c j.o/co-cm» f (SO»RO-«;I»PTI*((G?HP/3,U»»(RT"»3.r-pi»»-».j)n _ P? = J(!?O»RI)/('>O-PI))»(.(SI-SQ).-.(GPHO/2.0)»(RO»»2.C-RI»»2.:)) DO 15 I»l,N
PHI ( 11 =A J »R 11) -_( G 'HO/' . r ) »(R(I>»»3. ; | » PHI^dlsA.-CGRHO)» {R(I)»»?. j)
_ *>HIPP(I» = <-2.;)»(GPH0)'R(I) SIGR<n=PHI<I»/R(T)
. - SIGT(I)=oHIP(I) + GRM0»(R(I)»»2.a) 15 CONTINUE GC. TO >*■■_ 2L CONTINUE
THIS SECTION OF THF SUBROUTINE CONTAINS" THE CALCULATIONS .FOR A CONSTANT IN-^LAME SHEAR..STRESS.CRITERION
STR STR STR STR STP -SIR STR STR STR STR STR .SIR. STR STR STR STR STR STR STR STR STR STR STP 5TR. STR STo STR STR ST» STR STR STR ST? ST» STR STR_ STR STR STR STR STR 5TR STR STR STR STR STR
1 2 3 u 5
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IG 11
_12 13 l^ 15 16 17 ie 19 2C 21 22 23 2«« 25 26 27 29 29 j; 31 32 •» -i
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__ . .8TWfII«0HHl»/«tX1 STR 56 SIGT(T) = 0HI0(n*6oHnMR(I)"2.:) STR 57
_25 CONTINUE ' ' ST«? 5 <J GO TO *•! _STR-_59
JCCONUNUI — STR &: C THIS SECTION OF THE SUIRO'JTINE CONTAINS THE C»LCUL»TIO»« STP 61
- C FO^ A OlUuCE-MT CO-OITIOH AT THE INNER PAOIUS WITH A STR. S2 C STRESS CRITEPIOM OF CONSTANT THETA STRESS »JJ JJ
!'r,V^Ti--TAr/rR^rr; ♦on.-Po.Ro,i• ,,iT?ov;227T7^'p^(GrMoVA2FriTsTTR »« f(l.V3.:)»POMRO»'3.--RI^3..n) Tl? %j,
BC=S«»ROMG?MO/3. •)'(P'>,'3.])-Aj»PO STR ^^
DO 35 I = 1«N ... cTo 7- — nHi(T»=A-»R(n-»r,"MO/:.:»MR«n^3.:)*90 JTR ^ SHl»II»«\--.t5i».H0liW<X»t?l«iJ—. — — sT9" 72
PHippin = «-?.:>'«F,pMn)»p«Ti STO 73
SIGR(I)=owTn)/=!(T» - -— ST9 7u sicTm = pHi?m*GoH0M'(i»"2.")» STP 75
CONTINUE ST^ 7fc COMTfHUS STR _77 RETURN ~ ' STR 7« END
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