Addressing punching shear failure

7272019 Addressing punching shear failure

httpslidepdfcomreaderfulladdressing-punching-shear-failure 13 January 201314

design issues forstructural engineers

STRUCTURAL DESIGN

By Carlos E Ospina PhD PEand Neil M Hawkins PhDDist M ASCE Hon M ACI

Addressing Punching Failure

Considerations to Prevent

Premature Concentric

Punching Shear Failure in

Reinforced Concrete (RC)

Two-way Slabs

wo-way slabs are unique to ReinforcedConcrete (RC) construction Te mostcommon type due to its ease of form-ing and speed of construction is the

flat plate a slab of uniform thickness supported bycolumns without beams drop panels or capitalsFlat plates are common in building constructionand can also be found as deck components in

waterfront piers and wharves

Te design of RC flat plates is generally gov-erned by serviceability limits on deflection orby the punching shear capacity of the slab atthe slab-to-column interface or at locations ofconcentrated loads In practice for the specificcase where the transfer of unbalanced moment tothe column is minimal most punching failureslook alike a pyramid or truncated cone of slabremains around the column as the slab is loadedto failure Punching may occur before (brittle) orafter (ductile) a yield line mechanism has formedin the slab around the column Brittle punching is

undesirable because

there is little warn-ing of the impendingfailureFor about 50 years

ACI 318 has used theequation V n = 4radic f 983079c bod (where f 983079c is the specifiedcompressive strength of concrete bo is the criticalperimeter measured at 05d from the columnface and d is the effective depth of the flexuralreinforcement in the slab) for the nominal con-centric punching shear capacity of two-way RCslabs Tis expression was first introduced in the1963 code following recommendations provided

by ACI Committee 326 (Shear and Diagonalension) and is based on subtle modificationsto a design procedure developed by Moe (1961)Te ACI 318 equation (V ACI ) has served the

profession well However with the increasing useof higher strength steels and concretes the equa-tion is facing increasing scrutiny from researchersand practitioners because (1) it does not include afactor for the effect of the slab flexural reinforce-ment ratio ρ on the slab punching capacity and(2) average shear stresses significantly lower than4radic f 983079c at punching have been reported by severalresearchers for test slabs with ρ lt 1 and also for

slabs with d gt 8 inchesTis article discusses qualitatively the relevanceof these ldquoperceivedrdquo deficiencies in the ACI 318punching shear equation highlighting its short-comings and suggesting ways to improve theexisting code provisions Tis discussion concernsthe concentric punching shear capacity only anddoes not include the effects of transferring unbal-anced moments However the concepts suggestedhere can be readily extended to the moment trans-fer situation by use of the interaction relationshipdiscussed in R111172 of ACI 318

Reinforcement Ratio Effect

and Interaction Between

Shear and Flexure

Te absence of a ρ term is often cited as a majordeficiency by those who claim that the ACI 318equation does not predict the punching shearcapacities of test slabs as accurately as other equa-

tions that explicitly include this variable Tisclaim is however unfounded Alexander andHawkins (2005) reminded the profession thatthe ACI 318 equation was never intended to beused as a shear capacity predictor Instead it isa design equation aimed at precluding a brittlepunching shear failure before the slab developsits flexural capacity Its use assumes that the slabhas already been properly designed for flexureTe interaction between the transferred shearV

and the shear associated with the flexural capacityof the slab V flex as envisioned by Moe (1961) isqualitatively shown in Figure 1 Te unbroken

straight black line and the unbroken black curverepresent conditions for a flexural failure and ashear failure respectively When the designer pro-vides the proper amount of flexural reinforcementto resist the demand the flexural capacity matchesthe design load V flex plots as a straight line againstthe flexural load capacity because it is the productof the slab design load and the area tributary to thecolumn Point A represents the ldquobalanced failurerdquopoint ie the point where the slab fails simulta-neously in flexure and shear o the right of point

A V gt V flex ie shear failure governs o the leftof point A V flex lt V ie flexural failure governs

Te latter is the target failure zone for design-ers o allow full moment redistribution andthe development of sufficient slab deformationto warn of any impending failure Moe recom-mended that the slab be designed for V = 11V flex Tus the intersection of the steeper straightline with the shear design curve leads to a slightreduction in shear capacity (point B) Te pla-teau B-D is equivalent to the nominal punchingshear capacity V ACI In practical terms the designenvelope O-B-D separates flexural from shearfailures confirming that even though the ACI 318equation is not explicitly set up in terms of ρ it is

tacitly based on a term (V flex ) that accounts for ρ

Punching of Slabs with

Low Reinforcement Ratios

and the Issue of Ductility

Te effect of ρ on slab punching capacity has beendiscussed in the past by many researchers Intuitivelya decrease in ρ should lead to a reduction in thedepth of the compression zone available to resisttransverse shear and also to an increase in the width


httpslidepdfcomreaderfulladdressing-punching-shear-failure 23STRUCTURE magazine January 201315

of flexural cracks near the column Te increasein crack width should result in a reduction inaggregate interlock and dowel action Tecombination of those three effects should leadto a reduction in the punching shear capacity

Even though punching shear tests of slabs withlow ρ are few ndash the vast majority have beenperformed on slabs with fairly large amountsof flexural reinforcement to avoid flexural fail-ure ndash there is experimental evidence (Criswell1974 Guandalini et al 2009 and Widianto et al2009) indicating that two-way RC slabs with ρ

lt 1 may fail at shear stresses lower than 4radic f 983079c and display little ductility prior to punchingTe problem is exacerbated when d gt 8 inches(Guandalini et al 2009) With the use of higherstrength steels and higher strength concretes

many flat slabs now have ρ lt 1Te fact that a slab with low ρ can fail at a shear

stress less than 4radic f 983079c may seem to create the needfor a new equation for V n Such a ldquonecessityrdquo ishowever unjustified because as shown by Peirisand Ghali (2012) as long as the slab is properlydesigned for flexure the slab will reach V flex before

it punches Hence the maximum shear that canbe transferred by a slab with a low ρ is likely to bethat associated with the flexural capacity of theslab Tat shear can be calculated from a yieldline analysis assuming a concentric mechanism

centered on the column or concentrated load orif using a finite element program by extractingthe shear associated with the applied load on theslab at its flexural capacityHistorically the mode of failure of RC two-

way slabs tested in the laboratory has beendetermined by comparing the failure load V

Figure 1 Slab design rationale (After Moe (1961) and Alexander andHawkins (2005))

Figure 2 Relationship between V V ACI and V flex for slabs with varying ρ

O

B

B

C

D

A

FlexuralFailures

Design Load (Flexural Capacity)

T

r a n s f e r r e d

S h e a r

VFailures

Shear

V=Vflex

VACI

VMoe

V=11 Vflex

00

10

20

30

40

50

60

70

80

90

000 001 002 003 004 005 006 007

c f d b

V

0 c f d b

V

0 c f d b

V

0 c f d b

V

0 c f d b

V

0

ρ

ρ

ρ

θ

References

ACI-ASCE Committee 326 1962 ldquoShear and Diagonal ension Part 3ndashSlabs and Footingsrdquo ACI Journal March pp 353-395

ACI Committee 318 2011 Building Code Requirements for Structural Concrete (ACI 318-11) and Commentary (ACI 318R-11) AmericanConcrete Institute Farmington Hills MI

ACI-ASCE Committee 426 1974 ldquoTe Shear Strength of Reinforced Concrete Members ndash Slabsrdquo ASCE J of the Str Div V 100 NoS8 pp 1543-1591

Alexander SDB and Hawkins NM 2005 ldquoA Design Perspective on Punching Shearrdquo ACI Special Publication SP-232 Punching Shearin Reinforced Concrete Slabs pp 97-108

Criswell ME 1974 ldquoStatic and Dynamic Response of Reinforced Concrete Slab-Column Connectionsrdquo ACI Special Publication SP-42 Shear in Reinforced Concrete pp 721-746

Guandalini S Burdet O Muttoni A 2009 ldquoPunching tests of slabs with low reinforcement ratiosrdquo ACI Structural Journal V 106 No1 Jan-Feb pp 87-95

Moe J 1961 ldquoShearing Strength of Reinforced Concrete Slabs and Footings under Concentrated Loadsrdquo Development Department Bulletin

No D47 Portland Cement Association Skokie IL 130 ppMuttoni A 2008 ldquoPunching Shear Strength of Reinforced Concrete Slabs without ransverse Reinforcementrdquo ACI Structural Journal

V 105 No 4 July-Aug pp 440-450

Ospina CE Birkle G Widianto Wang Y Fernando SR Fernando S Catlin AC and Pujol S 2012 ldquoACI 445 Collected PunchingShear Databankrdquo httpsneesorgresources3660

Peiris C And Ghali A 2012 ldquo Flexural Reinforcement Essential for Punching Shear Resistance of Slabsrdquo SP 287-06 Recent Developmentsin Reinforced Concrete Slab Analysis Design and Serviceability American Concrete Institute Farmington Hills MI May

Widianto Bayrak O and Jirsa JO 2009 ldquowo-way Shear Strength of Slab-Column Connections Reexamination of ACI 318 Provisionsrdquo ACI Structural Journal V 106 No 2 March-April pp 160-170



against V flex with the latter determined from yield line analysis Brittleshear failure occurs if VV flex lt 1 whereas the failure is driven primarilyby flexure if VV flex gt 1 Additional refinements to account for strainhardening effects and in-plane restraint effects have been proposed toestablish the brittle versus ductile slab failure mode boundary Whetheran RC two-way slab falls into either category depends primarily on ρ

f 983079c and the geometric characteristics of the slab-column connectionIncreased ductility is expected in slabs with larger VV flex ratios but atthe expense of a punching capacity reduction For VV flex higher than

11 ndash ie low ρ ndash Moersquos theory assumes that such failures should bepreceded by significant deflection increases due to extensive yieldingof the slab reinforcement surrounding the columnUnfortunately direct comparisons between V and V flex do not neces-

sarily define whether a slab will deform significantly prior to punchingExperience shows that it is incorrect to assume that satisfying V = V ACI

will result in markedly increasing deflections or rotations at a slab-column connection before punching occurs Te reason is shown inFigure 2 where the shear force-rotation responses per the Critical ShearCrack Teory (CSC) of Muttoni (2008) are shown for three slabshaving identical geometries and differing ρ values Muttonirsquos theory isprobably the most accurate punching shear response predictor availableBlack solid circles represent punching failures per the CSC failure

criterion Red empty circles signal full yielding of the slab based on ayield line analysis and assuming a line of contra-flexure in the slab at022 times the distance between columns Te yield line capacity is thatfor a local failure mechanism centered on the column and is less thanthe mechanism associated with full yielding of the slab reinforcementTe slab with ρ = 03 was expected to reach V flex at a strength substan-

tially less than V ACI Punching failure in this case is driven by flexure andthe slab displayed considerable rotation before punching Te responseof the slab with ρ = 09 shows that even though V flex was expected tomatch V ACI punching occurred prematurely in a brittle fashion priorto developing the full local flexural capacity Tis result highlightsthe inadequacy of defining a slab failure mode using a strength-basedapproach only it implies that reaching V flex is not enough by itself to

prevent premature punching failure Te response of the slab with ρ =18 shows that this slab is expected to punch at a load similar to V ACI and well below V flex with no ductility whatsoever prior to punchingTe most significant limitation of using the VV flex approach to separate

shear and flexure-driven failures is that attention is concentrated on theload-resisting aspects of the slab response and not on the associateddeformations In 1963 the primary structural design emphasis was on theaccurate evaluation of strength with little attention paid to deformationslet alone the fact that none of the slabs examined by Moe corresponded tothe VV flex gt 10 case as noted by Widianto et al (2009) In fact excludingfootings none of the test slabs considered by Moe had ρ lt 1Te situation is even more serious for earthquake-resistant design

Even though concentric punching is linked mainly to gravity load-ing conditions the associated deformability issues can be invoked toattempt addressing those in the presence of lateral loads Experimentshave shown that the ductility under lateral loading increases as VV ACI decreases oday the basic concepts of seismic design including the needfor ductility and what ductility means are widely understood and usedFor performance to be acceptable in medium and high seismic designcategories (SDC) the flexural strength must be maintained throughdisplacements that are several times those at yielding of the flexuraltension reinforcement For one-way action there is always ductilitywhen the flexural strength is achieved prior to the shear strength Shearfailure following development of the flexural strength is only likely ifthe flexural reinforcement undergoes rapid strain hardening Even thenthe deformations have increased sufficiently that adequate warning has

been provided of the impending failure Unfortunately for two-wayaction in slabs deformations do not start to develop rapidly once theflexural strength is reached at the slab-column interface and the useof low reinforcement ratios in that region does not ensure ductility

Tese observations suggest supplementing the ACI 318 equation with adesign provision that explicitly addresses minimum deformability require-ments for RC two-way slabs to delay premature concentric punchingfailure One possible approach is a re-arrangement of Muttonirsquos CSCbased on a target ratio of slab rotation at ultimate to slab rotation at firstyield Once the ductility shortage is identified the most practical solutionis the addition of shear reinforcement However designers should never usea punching shear strength greater than that for the development of a localyield line mechanism centered on the column Guidance to evaluate V flex

for isolated slab-column connection tests and for slab systems is providedby the ACI-ASCE State-of-the-Art Report on punching of slabs (1974)

Size Effect

Another key consideration for reliably predicting the shear capacity oftwo-way slabs is the so-called size effect Figure 3 shows the effect ofincreasing the slab effective depth on the normalized punching capacity(4V test V ACI ) for selected test results extracted from the ACI 445 PunchingShear est Databank (Ospina et al 2012) Te shear capacity of two-wayslabs decreases as the effective depth increases Shear capacities less than4radic f 983079c develop for d gt 8 inches Te reduction in strength is substantialespecially if the slab is lightly reinforced Tese observations suggest theeffects of low reinforcement ratio and increasing slab depth are in largemeasure additive Both are detrimental to the shear capacity of the slabFigure 3 shows that reasonable punching shear capacity estimates resultfor slabs with d gt 8 inches if V ACI is multiplied by 3radicd (with d in inches)

Carlos E Ospina PhD PE is Senior Project Manager withBergerABAM Inc in Houston He is Co-Chair of ACI-ASCESubcommittee 445C (Punching Shear) and Leader of the TaskGroup responsible for the development of the Collected PunchingShear Test Result Databank He is a Fellow of ACI He can bereached at carlosospinaabamcom

Neil Hawkins PhD Dist M ASCE Hon M ACI is a ProfessorEmeritus of Civil and Environmental Engineering Universityof Illinois and Affiliate Professor of Civil and EnvironmentalEngineering University of Washington He may be reached atnmhawkinillinoisedu

Figure 3 Observed size effect on RC two-way slab punching capacity

00

10

20

30

40

50

60

70

80

90

100

0 5 10 15 20 25 30

co

test

f d b

V



of flexural cracks near the column Te increasein crack width should result in a reduction inaggregate interlock and dowel action Tecombination of those three effects should leadto a reduction in the punching shear capacity

Even though punching shear tests of slabs withlow ρ are few ndash the vast majority have beenperformed on slabs with fairly large amountsof flexural reinforcement to avoid flexural fail-ure ndash there is experimental evidence (Criswell1974 Guandalini et al 2009 and Widianto et al2009) indicating that two-way RC slabs with ρ

lt 1 may fail at shear stresses lower than 4radic f 983079c and display little ductility prior to punchingTe problem is exacerbated when d gt 8 inches(Guandalini et al 2009) With the use of higherstrength steels and higher strength concretes

many flat slabs now have ρ lt 1Te fact that a slab with low ρ can fail at a shear

stress less than 4radic f 983079c may seem to create the needfor a new equation for V n Such a ldquonecessityrdquo ishowever unjustified because as shown by Peirisand Ghali (2012) as long as the slab is properlydesigned for flexure the slab will reach V flex before

it punches Hence the maximum shear that canbe transferred by a slab with a low ρ is likely to bethat associated with the flexural capacity of theslab Tat shear can be calculated from a yieldline analysis assuming a concentric mechanism

centered on the column or concentrated load orif using a finite element program by extractingthe shear associated with the applied load on theslab at its flexural capacityHistorically the mode of failure of RC two-

way slabs tested in the laboratory has beendetermined by comparing the failure load V

Figure 1 Slab design rationale (After Moe (1961) and Alexander andHawkins (2005))

Figure 2 Relationship between V V ACI and V flex for slabs with varying ρ

O

B

B

C

D

A

FlexuralFailures

Design Load (Flexural Capacity)

T

r a n s f e r r e d

S h e a r

VFailures

Shear

V=Vflex

VACI

VMoe

V=11 Vflex

00

10

20

30

40

50

60

70

80

90

000 001 002 003 004 005 006 007

c f d b

V

0 c f d b

V

0 c f d b

V

0 c f d b

V

0 c f d b

V

0

ρ

ρ

ρ

θ

References

ACI-ASCE Committee 326 1962 ldquoShear and Diagonal ension Part 3ndashSlabs and Footingsrdquo ACI Journal March pp 353-395

ACI Committee 318 2011 Building Code Requirements for Structural Concrete (ACI 318-11) and Commentary (ACI 318R-11) AmericanConcrete Institute Farmington Hills MI

ACI-ASCE Committee 426 1974 ldquoTe Shear Strength of Reinforced Concrete Members ndash Slabsrdquo ASCE J of the Str Div V 100 NoS8 pp 1543-1591

Alexander SDB and Hawkins NM 2005 ldquoA Design Perspective on Punching Shearrdquo ACI Special Publication SP-232 Punching Shearin Reinforced Concrete Slabs pp 97-108

Criswell ME 1974 ldquoStatic and Dynamic Response of Reinforced Concrete Slab-Column Connectionsrdquo ACI Special Publication SP-42 Shear in Reinforced Concrete pp 721-746

Guandalini S Burdet O Muttoni A 2009 ldquoPunching tests of slabs with low reinforcement ratiosrdquo ACI Structural Journal V 106 No1 Jan-Feb pp 87-95

Moe J 1961 ldquoShearing Strength of Reinforced Concrete Slabs and Footings under Concentrated Loadsrdquo Development Department Bulletin

No D47 Portland Cement Association Skokie IL 130 ppMuttoni A 2008 ldquoPunching Shear Strength of Reinforced Concrete Slabs without ransverse Reinforcementrdquo ACI Structural Journal

V 105 No 4 July-Aug pp 440-450

Ospina CE Birkle G Widianto Wang Y Fernando SR Fernando S Catlin AC and Pujol S 2012 ldquoACI 445 Collected PunchingShear Databankrdquo httpsneesorgresources3660

Peiris C And Ghali A 2012 ldquo Flexural Reinforcement Essential for Punching Shear Resistance of Slabsrdquo SP 287-06 Recent Developmentsin Reinforced Concrete Slab Analysis Design and Serviceability American Concrete Institute Farmington Hills MI May

Widianto Bayrak O and Jirsa JO 2009 ldquowo-way Shear Strength of Slab-Column Connections Reexamination of ACI 318 Provisionsrdquo ACI Structural Journal V 106 No 2 March-April pp 160-170
















Size Effect





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40

50

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70

80

90

100

0 5 10 15 20 25 30

co

test

f d b

V
















Size Effect





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80

90

100

0 5 10 15 20 25 30

co

test

f d b

V

Documents

Addressing punching shear failure