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Geometry-Induced Rigidity in Nonspherical Pressurized Elastic Shells
A. Lazarus, H. C. B. Florijn, and P. M. Reis
*
EGS. Lab: Elasticity, Geometry and Statistics Laboratory, Department of Mechanical Engineering, Department of Civil &
Environmental Engineering, Massachusetts Institute of Technology, Cambridge, Massachusetts 02139, USA
(Received 16 July 2012; published 5 October 2012)
We present results from an experimental investigation of the indentation of nonspherical pressurized
elastic shells with a positive Gauss curvature. A predictive framework is proposed that rationalizes the
dependence of the local rigidity of an indented shell on the curvature in the neighborhood of the locus of
indentation, the in-out pressure differential, and the material properties. In our approach, we combine
classic theory for spherical shells with recent analytical developments for the pressurized case, and
proceed, for the most part, by analogy, guided by our own experiments. By way of example, our results
elucidate why an eggshell is significantly stiffer when compressed along its major axis, as compared to
doing so along its minor axis. The prominence of geometry in this class of problems points to the
relevance and applicability of our findings over a wide range of length scales.
DOI: 10.1103/PhysRevLett.109.144301 PACS numbers: 46.70.De, 81.70.Bt, 87.64.Dz
Shells are ubiquitous as both natural and engineered
structures, including viral capsids [1], pollen grains, col-
loidosomes [2], pharmaceutical capsules, exoskeletons,
mammalian skulls [3], pressure vessels, and architectural
domes. In addition to their aesthetically appealing form,
shells offer outstanding structural performance. As such,
they typically have the function of enclosure, containment,
and protection, with an eggshell being the archetypal ex-
ample [Fig. 1(a)]. The mechanics of thin elastic shells [4,5]
have long been known to be rooted in the purely geometric
isometric deformations of the underlying curved surface,
since stretching is energetically more costly than bending
[6]. However, despite a vast literature on the isometry of
rigid surfaces [7,8], establishing a general direct theoretical
connection between the differential geometry of surfaces
and the mechanics of shells is a challenging endeavor. This
is due to the difficulties in systematically quantifying the
effect of a small but finite thickness on the mechanical
response of a curved surface. To circumnavigate this issue,
explicit boundary value problem calculations tend to be
performed by deriving the local equilibrium equations of
3D continuum mechanics and then taking the limit of a
small but finite thickness within the kinematics [9]. For
example, closed analytical solutions have been obtained in
this fashion for the indentation of spherical unpressurized
[10,11] and pressurized [12] shells. For more intricate
geometries and mechanical environments, numerical
methods can be used such as full scale finite element
simulations [13,14]. Although powerful, these computa-
tional approaches can sometimes come at the detriment
of physical insight and predictive understanding of the
interplay between the mechanics of the structure and the
geometry of its surface.
Here, we study the effect of the geometry of surfaces
with a positive Gauss curvature on the linear mechanical
response under the indentation of thin elastic shells, with or
without an in-out pressure differential. Our goal is to
quantify the geometry-induced rigidity (GIR) which we
define as the amount by which a nonspherical shell is
stiffened when compared to a spherical shell with the
same thickness and material properties [15]. For this pur-
pose, we perform precision desktop-scale experiments
where both the geometry of the shells and their material
properties are accurately custom controlled using rapid
prototyping and digital fabrication techniques. The rigidity
is quantified through indentation tests and the differential
pressure is set by a syringe-pump system under feedback
control. First, we study the rigidity at the pole of ellipsoidal
shells with different aspect ratios toward investigating the
0 0.5 1 1.5 2
0
100
200
300
400
500
F (mN)
δ (mm)
0 0.5 1 1.5 2
0
20
40
60
F (mN)
δ (mm)
F
F
F
F
b/a = 2 b/a = 1.5
b/a = 1 b/a = 0.5
b/a
K
(b)
K
p
(a)
Egg
(c)
b
a
δ
FIG. 1 (color online). (a) A chicken’s egg and 500 m thick
ellipsoidal shell samples with four aspect ratios b=a ¼
f2; 1:5; 1; 0:5g and four elastomers with Young’s modulus E ¼
f0:2; 0:5; 0:6; 1g MPa. (b) Load-displacement curves for inden-
tation at the pole for an in-out differential pressure p ¼ 0Pa,
E ¼ 1 MPa, and three different aspect ratios b=a ¼ 0:5, 1, and
2. (c) Load-displacement curves for indentation at the pole for
b=a ¼ 1:5, E ¼ 1 MPa, and 0 <p<10 kPa.
PRL 109, 144301 (2012)
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specific case of axisymmetric convex surfaces. We then
turn to ellipsoids indented along their meridian, where the
surface has two distinct principal curvatures. Combining
our experimental results with the classic theory of elastic-
ity of spherical shells due to Reissner [10] as well as more
recent developments by Vella et al. [12], we find that the
effective rigidity is induced by (i) the local curvature in the
neighborhood of the locus of indentation, (ii) the material
properties of the shell, and (iii) the in-out differential
pressure.
Fabricating our thin shells involves the conception of a
computer assisted design model for a mold of the target
shell structure, which is then 3D printed out of ABS plastic,
and used to cast samples with Vinylpolysiloxane (a silicone-
based elastomer). In our experiments, we use elastomers
with four values of the Young’s modulus in the range
0:2 <E<1MPaand a Poisson ratio of 0:5.Inspired
by the eggshell geometry [Fig. 1(a )], we print molds wit h
four ellipsoidal shapes that set the geometric parameters to
be four polar radii, b ¼f1: 25; 2:5; 3:75; 5g cm, an equato-
rial circular radius, a ¼ 2:5cm, and a thicknes s t ¼
0:5mm(the latter two parameters are set constant for all
experiments). This value of thickness (t=a ¼ 0:02 1)en-
sures the validity of the thin shell assumption. In Fig. 1(a),we
show photographs of four representati v e sample s made of
different polymers (different colors correspond to different
values of E) with aspect ratios b=a ¼f0:5; 1; 1:5; 2g.
We start by vertically indenting the pole of unpressur-
ized shell specimens using a hemispherical cap indentor
(1.5 mm radius) at the constant speed of 1mm= min ,
ensuring point indentation under quasistatic conditions.
The compressive force F resulting from the indentation
by the imposed displacement is recorded using the load
cell of an Instron machine with a resolution of 100 N.
Typical load-displacement curves for shells with different
aspect ratios are presented in Fig. 1(b). We focus on the
first linear regime occurring for =t < 1 and define the
experimental rigidity K as the initial slope in the load-
displacement curves. For a given material and shell thick-
ness, the rigidity increases with the aspect ratio, suggesting
the possibility of enhancing the rigidity of a thin elastic
shell by simply changing its geometry.
To investigate this GIR behavior observed in Fig. 1(b),
we now measure K at the pole for 16 unpressurized shell
specimens (four aspect ratios b=a and four values of E).
We quantify the GIR as the dimensionless ratio K=K
s
where K
s
is the theoretical rigidity of unpressurized spheri-
cal shells under indentation defined by Reissner [10]as
K
s
¼ 8D=L
2
b
. Here, D ¼ Et
3
=12ð1
2
Þ is the bending
modulus of the shell, L
b
¼ðDa
2
=EtÞ
1=4
is a characteristic
length scale arising from the balance of bending and
stretching, and a is the radius chosen for reference. In
Fig. 2(b), we present the dependence of the GIR at the
pole on the shell’s aspect ratio, finding that it scales pro-
portional to b=a; i.e., for a given material, an ellipsoidal
shell that is twice as high as the spherical counterpart is
also twice as rigid.
We rationalize the above observation by considering the
equations for static equilibrium of thin elastic shells of
revolution under indentation [5] and computing the me-
chanical stresses at the pole. In its stress-free reference
configuration, the shell can be represented by the generat-
ing planar curves ¼ ðsÞ and z ¼ zðsÞ, where s is the
arc-length along the shell’s midplane [see the schematic
diagram in Fig. 2(a)]. In order to understand the linear
mechanical response, we consider the particular case of
small strains and small rotations under indentation at the
convex pole. Assuming axisymmetric and twistless defor-
mations, the linear equilibrium equations can be written in
the radial and axial directions, in term of s alone, as
d
ds
ð
s
0
Þ
¼ 0; (1)
1
d
ds
ð
s
z
0
Þ
D
d
ds
ð
0
Þ
0
¼f
z
; (2)
where ðsÞ is the rotation angle between the reference and
deformed configuration,
s
ðsÞ and
ðsÞ are the axial and
hoop stresses, respectively, f
z
ðsÞ is the force density per
unit area in the axial direction, and ðÞ
0
denotes differentia-
tion with respect to s. For our particular ellipsoidal geome-
tries, the generating planar curve is parametrized by
zðsÞ¼b
ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi
1
2
ðsÞ=a
2
p
. In the neighborhood of the inden-
tation point at the flat pole of the ellipsoid, where the
deformations occur, the local generating plane curve can
be described by the linear approximations ðsÞs,
10
−2
10
0
10
2
10
1
10
2
τ
K / (κ
2
EtD)
1/2
Eq. (4)
b/a = 0.5
b/a = 1
b/a = 1.5
b/a = 2
0 1000 2000 3000 4000
1
2
3
4
5
6
p (Pa)
GIR, K/Ks
Eq. (4)
b/a = 0.5
b/a = 1
b/a = 1.5
b/a = 2
0 0.5 1 1.5 2 2.5
0
0.5
1
1.5
2
2.5
b/a
GIR, K/Ks
Eq. (3)
E = 1 Mpa
E = 0.6 MPa
E = 0.5 MPa
E = 0.2 MPa
(a)
(b)
(c)
(d)
a
b
F
p
θ
ρ
δ
z
s
FIG. 2 (color online). Indentation at the pole. (a) Schematic of
a deformed ellipsoidal specimen. (b) Dependence of GIR K=K
s
on b=a for four different elastic polymers and p ¼ 0. (c) GIR
K=K
s
, as a function of the internal pressure p for four aspect
ratios b=a ¼f0:5; 1; 1:5; 2g and E ¼ 1 MPa. (d) Dimensionless
experimental rigidity Kl
2
b
=D against dimensionless pressure .
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0
ðsÞ1, and z
0
ðsÞs=r, where r ¼ a
2
=b is the local
radius of curvature. For s r, the angle ðsÞþz
0
ðsÞ
[Fig. 2(a)] which measures the direction of the tangent
along a deformed meridian, must vanish since this tangent
is perpendicular to the pushing vertical indentor and
ðsÞ¼z
0
ðsÞ¼s=r. Replacing this local kinematics
into the equilibrium equations (1) and (2), we find that
the membrane stresses are given by ðsÞ Z EðsÞ and
therefore proportional to the local curvature at the pole,
¼ 1=r ¼ b=a
2
(the two principal curvatures are equal
due to axisymmetry, ¼
1
¼
2
). After substituting the
global natural bending length scale L
b
introduced earlier
by its local counterpart, l
b
¼ðD=Et
2
Þ
1=4
, into K
s
, we can
generalize Reissner’s rigidity at the pole of nonspherical
thin elastic shells of revolution:
K
0
t
¼
4Et
2
ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi
3ð1
2
Þ
p
: (3)
In particular, for an ellipsoidal shell where ¼ b=a
2
,we
find that K
0
t
¼ðb=aÞK
s
, which is in excellent agreement
with the experimental GIR data as a function of aspect
ratios b=a presented in Fig. 2(b).
Having rationalized our results for unpressurized shells,
we now investigate the influence of an in-out differential
pressure on the rigidity of nonspherical axisymmetric
shells. We focus on four elastomeric ellipsoidal shells with
aspect ratios b=a ¼f0:5; 1; 1:5; 2g and E ¼ 1MPa,and
measure the rigidity K at their pole for differential pressures
in the range 0 <p<4kPa. Typical load-displacement
curves for increasing p are shown in Fig. 1(c) for a shell
with b=a ¼ 1:5. The corresponding dependence of the
GIR (K=K
s
), on p is presented in Fig. 2(c) and we find
that both pressure and geometry can enhance the rigidity.
Above, we showed for the unpressurized case that the
coupling between geometry and the mechanical response
was dictated by the local curvature. This now motivates us
to proceed by hypothesizing that the local nature of the
GIR also translates into the pressurized case. If this
hypothesis is correct, we should therefore find that the
indentation at the pole of a pressurized ellipsoidal shell is
locally identical to the indentation of a pressurized spheri-
cal shell of radius r ¼ 1 = ¼ a
2
=b. As such, we seek a
generalization of the theoretical rigidity of the pressurized
spherical shell of Vella et al. [12],
K
p
t
¼
8D
l
2
b
ð
2
1Þ
1=2
2 arctanhð1
2
Þ
1=2
; (4)
where we replace a by the local radius of the curvature
at the pole r ¼ 1=, where ¼
1
¼
2
, so that ¼
pr
2
=4ðEDtÞ
1=2
is a local dimensionless internal pressure
and l
b
is the local bending scale at the pole defined above.
The dependence of the dimensionless experimental rigid-
ity, Kl
2
b
=D, on the local dimensionless pressure is
showed in Fig. 2(d), where we find that all the experimental
data collapse onto the predicted rigidity K
p
t
(dashed line).
This excellent agreement confirms that Eq. (4) accurately
captures the experimental rigidity at the pole of pressurized
thin ellipsoidal shells, which corroborates our hypothesis.
We progress by generalizing the current framework to
nonaxisymmetric convex shells which are locally charac-
terized by two distinct principal curvatures
1
and
2
,a
Gauss curvature
G
¼
1
2
, which we restrict to be posi-
tive, and a mean curvature
M
¼ð
1
þ
2
Þ=2 [Fig. 3(a)].
For the moderately elongated ellipsoids we consider in our
study,
M
ffiffiffiffiffiffi
G
p
, as shown in Fig. 3(b), with representa-
tive schematics in the inset. Encouraged by our previous
success in characterizing the rigidity of convex axisym-
metric shells, we proceed again by analogy and hypothe-
size that the local mean curvature
M
generalizes the role
of in the GIR of nonaxisymmetric shell surfaces. Note
that, as required, this generalization reduces to
M
¼
1
¼
2
¼ at the pole. Thus, replacing ¼ 1 =r by
M
in
Eqs. (3) and (4), we propose a combined expression that
describes the rigidity of unpressurized or pressurized non-
spherical shells with a positive Gauss curvature,
K
t
¼
4Et
2
ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi
3ð1
2
Þ
p
M
=2ð
2
1Þ
1=2
arctanhð1
2
Þ
1=2
; (5)
where is the local dimensionless pressure introduced
above with r ¼ 1=
M
.
1/
1
1/
2
F
10
0
10
2
10
1
10
2
τ
K / (κ
M
2
EtD)
−1/2
Eq. (5)
κ
M
=26 m
−1
κ
M
=32 m
−1
κ
M
=39 m
−1
κ
M
=57 m
−1
0 50 100
0
50
100
κ
G
(s)
1/2
(m
−1
)
κ
M
(s) (m
−1
)
κ
G
= κ
M
2
b/a = 1.5
b/a = 2
0 50 100
0
50
100
GIR, K(s)/K
s
(s)
κ
M
(s) (m
−1
)
Eq. (5), τ = 0
b/a = 1.5
b/a = 2
(a)
(c)
F
A
(d)
(b)
s
2
1
B
A
B
FIG. 3 (color online). (a) Experimental setup for shells with
two distinct principal curvatures
1
and
2
. (b) Evolution of the
Gauss and mean curvatures along the meridian of the ellipsoids
with b=a ¼ 1:5 and 2. (c) Evolution of geometry-induced rigid-
ity K=K
s
ðsÞ with mean curvature
M
ðsÞ for two unpressurized
shells with aspect ratios b=a ¼ 1:5; 2 and E ¼ 1 MPa.
(d) Evolution of dimensionless experimental rigidity
KðsÞl
2
b
ðsÞ=DðsÞ with dimensionless pressure for four points
along the meridian of a pressurized shell with b=a ¼ 2 and
E ¼ 1 MPa.
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A modification of our experimental setup that allows for
indentation in regions with two different principal curva-
tures,
1
and
2
, is now used to test the proposed Eq. (5).
For this, two shells with Young’s modulus E ¼ 1 MPa
and aspect ratios b=a ¼ 1:5 and 2 are marked every
s
i
¼ 0:5cm along one meridian from the pole to the
base [Fig. 3(a)]. Once pressurized to a set constant p, the
shells are indented at the points Pðs
i
Þ with the compressive
force F aligned perpendicularly to the surface. For a given
differential pressure p, we measure the experimental local
mean curvature
M
ðs
i
Þ using digital image processing at
each indentation point Pðs
i
Þalong the meridian of our shell
specimen [Fig. 3(b)].
In Fig. 3(c) , we plot the dependence of the Geometry-
Induced Rigidity, Kðs
i
Þ=K
s
ðs
i
Þ, on the mean curvature
M
ðs
i
Þ for two unpressurized shells with aspect ratios
b=a ¼ 1:5 and 2 [Reissner’s result per unit radius K
s
ðs
i
Þ¼
4Etðs
i
Þ
2
=
ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi
3ð1
2
Þ
p
is used as the reference rigidity
K
s
ðs
i
Þ]. As expected, the measured Kðs
i
Þ scales linearly
with Reissner’s rigidity K
s
ðs
i
Þ with a prefactor
M
ðs
i
Þ, i.e.,
Kðs
i
Þ¼
M
ðs
i
ÞK
s
ðs
i
Þ, which conforms to Eq. (5) with a
dimensionless differential pressure ¼ 0. The role of
M
in moderately nonaxisymmetric shells also extends to the
pressurized case, as shown in Fig. 3(d), where we plot the
evolution of dimensionless experimental rigidities
KðsÞl
b
ðsÞ
2
=D as a function of dimensionless pressure
ðsÞ for four indentation points Pðs
i
Þ along the shell with
b=a ¼ 2. We find excellent agreement between the experi-
mental rigidities Kðs
i
Þ and the predicted K
t
ðs
i
Þ (dashed
line) confirming the accuracy of Eq. (5).
Having learned about the geometry-induced rigidity in
convex elastic shells through Eq. (5), we turn our learnings
into an inverse problem toward developing a nondestruc-
tive method to measure the parameters of the shell
(e.g., shell thickness) upon knowing the geometry of the
underlying surface and the local mechanical response.
Fortunately, the thickness profile of our experimental
samples varied slightly along the meridian of our shells
due to limitations of our fabrication procedure; the male
and female parts of the molds were not perfectly aligned
during casting. Earlier, when we tested Eqs. (3)–(5), we
took for the thickness, tðs
i
Þ, the local average measured
over a curvilinear distance s ¼ 3 l
b
in the neighborhood
of the indentation point Pðs
i
Þ obtained by cutting the shell
in half and analyzing the profile with a digital flat scanner
[Fig. 4 (inset)]. Note that in our experiments, the local
bending scale, 1:6 <l
b
¼ðD=Et
2
M
Þ
1=4
< 2:5mm,isof
the order of the diameter of the hemispherical cap indentor
¼ 3mm. In Fig. 4, we compare this average thickness
tðs
i
Þ along the meridian of two shell specimens measured
directly from the scanned profiles of the cut shells, with the
thickness predicted from Eq. (5) with ¼ 0, through
indentation. The variation of the thickness profile is accu-
rately captured within a spatial resolution of s 2l
b
5mm, confirming the relevance of the local bending scale
l
b
as a measure of locality. More importantly, the fact that
we are able to recover the thickness profile using Eq. (5)
demonstrates that our proposed description can be used as
a precision nondestructive technique for nonspherical thin
elastic shell with a positive Gauss curvature.
In summary, we have quantified the relation between
geometry and the mechanical rigidity of positively curved
thin elastic shells under pressure, arriving to our predictive
framework primarily by analogy and an analysis of preci-
sion experimental data. Our work calls for a formal iden-
tification, in particular, for a higher degree of asphericity
¼ð
2
1
Þ=ð
2
þ
1
Þ, where our description is likely
to be challenged, which will require a theoretical effort that
we hope our work will help catalyze. Still, the range of
curvatures we have considered (0 <<0:6 for which
M
ffiffiffiffiffiffi
G
p
), is applicable and relevant across a wide range
of practical instances of natural and man-made nonspher-
ical shells. The scale invariance of geometry-induced
rigidity suggests that our framework should find uses
across length scales: from the mechanical testing of viral
capsids through atomic force microscopy [16], to ocular
tonometry procedures [17] or in the design of architectural
shells [18]. Moreover, one can now interpret the increased
stiffness of a chicken egg compressed along its poles as
compared to doing so along the equator as being attributed
to geometry-induced rigidity.
We thank C. Perdigou for help with preliminary experi-
ments and B. Audoly, D. Vella, A. Vaziri, and A. Boudaoud
for enlightening discussions.
Note added.—During the preparation of our manuscript,
we became aware of the complementary theoretical and
numerical work of D. Vella et al. [19], on a similar topic.
0 1 2 3 4 5
0
0.2
0.4
0.6
0.8
Curvilinear distance, s (cm)
Thickness, t (mm)
theory
b/a = 1.5
b/a = 2
3.75 cm
2.5 cm
t
s
FIG. 4 (color online). Comparison between measured and pre-
dicted thickness profile for two unpressurized shells with b=a ¼
1:5 and 2 and 0:9 <E<1 MPa. Solid and dashed lines repre-
sent mean value standard deviation for the predicted thick-
ness. (Inset) Scanned image of a shell with aspect ratio
b=a ¼ 1:5 cut in half.
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*To whom correspondence should be addressed.
preis@mit.edu
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[15] Note that our notion of GIR in thin shells is not to be
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