土木工程外文翻译-原文

1、外文原文Response of a reinforced concrete infilled-frame structure to removal of two adjacent columnsMehrdad Sasan_Northeastern Uni versity, 400 Sn ell Engin eeri ng Cen ter, Bosto n, MA 02115, Un itedStatesReceived 27 June 2007; received in revised form 26 December 2007; accepted 24 Jan uary 2008Availa

2、ble on li ne 19 March 2008AbstractThe response of Hotel San Diego, a six-story reinforced concrete infilled-frame structure, is evaluated followi ng the simulta neous removal of two adjace nt exterior colu mns. An alytical models of the structure using the Finite Eleme nt Method as well as the Appli

3、ed Eleme nt Method are used to calculate global and local deformati ons. The analytical results show good agreement with experimental data. The structure resisted progressive collapse with a measured maximum vertical displaceme nt of only one quarter of an inch (6.4 mm). Deformation propagation over

4、 the height of the structure and the dynamic load redistribution following the column removal are experime ntally and an alytically evaluated and described. The differe nee betwee n axial and flexural wave propagati ons is discussed. Three-dime nsional Viere ndeel (frame) actio n of the tran sverse

5、and Ion gitud inal frames with the participati on of in fill walls is identified as the major mechanism for redistribution of loads in the structure. The effects of two potential brittle modes of failure (fracture of beam sections without ten sile rein forceme nt and reinforcing bar pull out) are de

6、scribed. The resp onse of the structure due to additi onal gravity loads and in the abse nee of in fill walls is an alytically evaluated.c 2008 Elsevier Ltd. All rights reserved.Keywords: Progressive collapse; Load redistributi on; Load resista nee; Dyn amic resp on se; Non li near an alysis; Brittl

7、e failure1. IntroductionThe prin cipal scope of specificati ons is to provide gen eral prin ciples and computati on al methods in order to verify safety of structures. The “ safety factor ” , which accor ding to moder n trends is in depe ndent of the n ature and comb in ati on of the materials u sed

8、, can usually be defi ned as the ratio betwee n the con diti ons. This ratio is also prop orti onal to the inv erse of the probability ( risk ) of failure of the structure.Failure has to be con sidered n ot only as overall collapse of the structure but also as un serviceability or, accordi ng to a m

9、ore precise. Common defi niti on. As the reach ing of a “limit state ” which causes the construction not to accomplish the task it was desi gned for. There are two categories of limit state :(1) Ultimate limit sate, which corresp onds to the highest value of the load-beari ng cap acity. Examples in

10、clude local buckli ng or global in stability of the structure; failure of some secti ons and subseque nt tran sformatio n of the structure into a mecha ni sm; failur e by fatigue; elastic or plastic deformati on or creep that cause a substa ntial cha nge of t he geometry of the structure; and sen si

11、tivity of the structure to alter nat ing loads, to fir e and to explosi ons.(2) Service limit states, which are fun cti ons of the use and durability of the structure. E xamples in clude excessive deformati ons and displaceme nts without in stability; early o r excessive cracks; large vibrati ons; a

12、nd corrosi on.Computati onal methods used to verify structures with respect to the differe nt safety co n diti ons can be separated in to:(1) Determi nistic methods, in which the ma in parameters are con sidered as nonran dom parameters.(2) Probabilistic methods, in which the mai n parameters are co

13、n sidered as ran dom para meters.Alternatively, with respect to the differe nt use of factors of safety, computati onal meth ods can be separated in to:(Allowable stress method, in which the stresses computed un der maximum loads ar e compared with the stre ngth of the material reduced by give n saf

14、ety factors.(2) Limit states method, in which the structure may be proporti oned on the basis of its maximum stre ngth. This stre ngth, as determ ined by rati onal an alysis, shall n ot be less tha n that required to support a factored load equal to the sum of the factored live load and dead load (

15、ultimate state ).The stresses corresp onding to work ing ( service ) con diti ons with un factored live and dead loads are compared with prescribed values ( service limit state ) . From the four possible comb in ati ons of the first two and sec ond two methods, we can obta in some u seful computati

16、onal methods. Gen erally, two comb in ati ons prevail:(1)determi nistic methods, which make use of allowable stresses. (2)Probabilistic meth ods, which make use of limit states.The main adva ntage of probabilistic approaches is that, at least in theory, it is possible to scie ntifically take into ac

17、co unt all ran dom factors of safety, which are the n comb in ed to defi ne the safety factor. probabilistic approaches depe nd upon :(1) Ran dom distributio n of stre ngth of materials with respect to the con diti ons of fabri catio n and erecti on ( scatter of the values of mecha ni cal properties

18、 through out the str ucture ); (2) Un certa inty of the geometry of the cross-secti on sand of the structure ( fa ults and imperfect ions due to fabricati on and erect ion of the structure );(3) Un certa inty of the predicted live loads and dead loads act ing on the structure; (4)U n certa inty rela

19、ted to the approximati on of the computati onal method used ( deviati on of the actual stresses from computed stresses ). Furthermore, probabilistic theories me an that the allowable risk can be based on several factors, such as :(1) Importa nee of the con struct ion and gravity of the damage by its

20、 failure; (2)Numbe r of huma n lives which can be threate ned by this failure; (3)Possibility an d/or likeliho od of repairi ng the structure; (4) Predicted life of the structure. All these factors are rel ated to econo mic and social con siderati ons such as(1) In itial cost of the con structi on;(

21、2) Amortizati on funds for the durati on of the con structi on;(3) Cost of physical and material damage due to the failure of the con struct ion;(4) Adverse impact on society;(5) Moral and psychological views.The defi niti on of all these parameters, for a give n safety factor, allows con structio n

22、 at the optimum cost. However, the difficulty of carry ing out a complete probabilistic an alysis has to be take n into acco unt. For such an an alysis the laws of the distributio n of the live load and its in duced stresses, of the scatter of mecha ni cal properties of mat erials, and of the geomet

23、ry of the cross-sect ions and the structure have to be known. F urthermore, it is difficult to in terpret the in teracti on betwee n the law of distributio n of stre ngth and that of stresses because both depe nd upon the n ature of the material, on t he cross-sect ions and upon the load act ing on

24、the structure. These practical difficulties can be overcome in two ways. The first is to apply differe nt safety factors to the mate rial and to the loads, without n ecessarily adopt ing the probabilistic criteri on. The seco nd is an approximate probabilistic method which in troduces some simplify

25、ing assump tions ( semi-probabilistic methods ) . As part of mitigation programs to reduce the likelihood of mass casualties following local damage in structures, the General Services Admi nistrati on 1 and the Departme nt of Defense 2 developed regulatio ns to evaluate progressive collapse resista

26、nee of structures. ASCE/SEI 7 3 defi nes progressive collapse as the spread of an in itial local failure from eleme nt to eleme nt eve ntually result ing in collapse of an en tire structure or a disproportio nately large part of it. Following the approaches proposed by Ellinwood and Leyendecker 4, A

27、SCE/SEI 7 3 defines two general methods for structural design of buildings to mitigate damage due to progressive collapse: in direct and direct desig n methods. Gen eral build ing codes and sta ndards 3,5 use in direct desig n by in creas ing overall in tegrity of structures. I ndirect desig n is al

28、so used in DOD 2. Although the in direct design method can reduce the risk of progressive collapse 6,7 estimation of post-failure performa nee of structures desig ned based on such a method is not readily possible. One approach based on direct design methods to evaluate progressive collapse of struc

29、tures is to study the effects of in sta ntan eous removal of load-beari ng eleme nts, such as colu mns. GSA 1 and DOD 2 regulati ons require removal of one load bearing element. These regulations are meant to evaluate general integrity of structures and their capacity of redistributing the loads fol

30、lowing severe damage to on ly one eleme nt. While such an approach provides in sight as to the exte nt to which the structures are susceptible to progressive collapse, in reality, the in itial damage can affect more than just one column. In this study, using analytical results that are verified aga

31、inst experime ntal data, the progressive collapse resista nee of the Hotel San Diego is evaluated, following the simultaneous explosion (sudden removal) of two adjace nt colu mns, one of which was a corner colu mn. In order to explode the colu mns, explosives were in serted in to predrilled holes in

32、 the colu mns. The colu mns were the n well wrapped with a few layers of protective materials. Therefore, n either air blast nor flyi ng fragme nts affected the structure.Fig, 1 A souih view hole! San Diegn. Cenr strucluiu is studied in this pape 匚F唱 2. Second floor of budding i Looking south L2. Bu

33、ildi ng characteristicsHotel San Diego was con structed in 1914 with a south annex added in 1924. The annex in cluded two separate build in gs. Fig. 1 shows a south view of the hotel. Note that in the picture, the first and third stories of the hotel are covered with black fabric. The six story hote

34、l had a non-ductile rein forced con crete (RC) frame structure with hollow clay tile exterior in fill walls. The in fills in the annex con sisted of two withes (layers) of clay tiles with a total thick ness of about 8 in (203 mm). The height of the first floor was about 190 -800 (6.00 m). The height

35、 of other floors and that of the top floor were 100 -600 (3.20 m) and 160 -000 (5.13 m), respectively. Fig. 2 shows the sec ond floor of one of the annex buildi ngs.Fig. 3 shows a typical pla n of this build ing, whose resp onse followi ng the simulta neous removal (explosi on) of colu mns A2 and A3

36、 in the first (ground) floor is evaluated in this paper. The floor system consisted of on e-way joists running in the Ion gitud inal directi on (North -South), as show n in Fig. 3. Based on compression tests of two concrete samples, the average concrete compressive strength was estimated at about 45

37、00 psi (31 MPa) for a standard con crete cyli nder. The modulus of elasticity of con crete was estimated at 3820 ksi (26 300 MPa) 5. Also, based on tension tests of two steel samples having 1/2 in (12.7 mm) square sections, the yield and ultimate tensile strengths were found to be 62 ksi (427 MPa) a

38、nd 87 ksi (600 MPa), respectively. The steel ultimate ten sile strain was measured at 0.17. The modulus of elasticity of steel was set equal to 29 000 ksi (200000 MPa). The build ing was scheduled to be demolished by implosi on. As part of the demoliti on process, the in fill walls were removed from

39、 the first and third floors. There was no live load in the buildi ng. All non structural eleme nts in cludi ng partiti ons, plumb ing, and furn iture were removed prior to implosi on. Only beams, colu mn s, joist floor and in fill walls on the peripheral beams were prese nt.3. Sen sorsCon crete and

40、steel strain gages were used to measure cha nges in strai ns of beams and colu mns. Lin ear pote ntiometers were used to measure global and local deformatio ns. The con crete strain gages were 3.5 in (90 mm) long hav ing a maximum strain limit of 0.02.壬he steel strain gages could measure up to a str

41、ain of0.20. Thestra in gages could operate up to a several hun dred kHz sampli ng rate. The sampli ng rate used in the experiment was 1000 Hz. Potentiometers were used to capture rotation (integral of curvature over a length) of the beam end regions and global displaceme nt in the build ing, as desc

42、ribed later. The pote ntiometers had a resoluti on of about 0.0004 in (0.01 mm) and a maximum operati onal speed of about 40 in/s (1.0 m/s), while the maximum recorded speed in the experime nt was about 14 in/s (0.35m/s).Fig, 3. Typical plan ol Hold San Diego (South Annex k Firsi door ivmoved column

43、s arc crosscii.:A1 :2Q rwm)3 口 W(1D mm)Tffiieal丄 IM I I144h4 3皆,0 mm) 20G mm|iA2：r-mm2 耿81 托 mm5&6 mmgas/rfiemm銘&P蝕知障PITCH DO mnrri SP 四 57mm) TypcailIig. 4. Rtmforvmcnt dbuiil of cokmnm ;ind (ai)险 am A3-B3 iiri corid flwr;乱 ndl (b) Bcsiin Al-A 2.4. Fin ite eleme nt modelUsing the finite element met

44、hod (FEM), a model of the building was developed in the SAP2000 8 computer program. The beams and columns are modeled with Berno ulli beam eleme nts. Beams have T or L sect ions with effective flange width on each side of the web equal to four times the slab thickness 5. Plastic hinges are assignedt

45、o all possible locations where steel bar yielding can occur, including the ends of elements as well as the reinforcing bar cut-off and bend locations. The characteristics of the plastic hin ges are obta ined using sect ion an alyses of the beams and colu mns and assu ming a plastic hinge len gth equ

46、al to half of the secti on depth. The current version of SAP2000 8 is not able to track formation of cracks in the elements. In order to find the proper flexural stiffness of sections, an iterative procedure is used as follows. First, the build ing is an alyzed assu ming all eleme nts are un cracked

47、. Then, mome nt dema nds in the eleme nts are compared with their crack ing bending moments, Mcr . The moment of inertia of beam and slab segments are reduced by a coefficie nt of 0.35 5, where the dema nd exceeds the Mcr. The exterior beam cracking bending moments under negative and positive moment

48、s, are 516 k in (58.2 kN m) and 336 k in (37.9 kN m), respectively. Note that no cracks were formed in the columns. Then the building is reanalyzed and moment diagrams are re-evaluated. This procedure is repeated until all of the cracked regions are properly identified and modeled.The beams in the b

49、uilding did not have top reinforcing bars except at the end regi ons (seeFig. 4). For in sta nee, no top rein forceme nt was provided bey ond the bend in beam A1 -2, 12 inches away from the face of column A1 (see Figs. 4 and 5). To model the pote ntial loss of flexural stre ngth in those secti on s,

50、 localized crack hin ges were assigned at the critical locations where no top rebar was present. Flexural stre ngths of the hin ges were set equal to Mcr. Such secti ons were assumed to lose their flexural stre ngth whe n the imposed bending mome nts reached Mcr.Fig. 5. Location of bends in beam top

51、 reinforcement fin an adjacent annex building at a location similar to beam A1-A2. close to column A ).The floor system con sisted of joists in the Ion gitud inal direct ion (North -South). Fig. 6 shows the cross section of a typical floor. In order to account for potential non li near resp onse of

52、slabs and joists, floors are molded by beam eleme nts. Joists are modeled with T-sect ions, hav ing effective flange width on each side of the web equal to four times the slab thick ness 5. Give n the large joist spaci ng betwee n axes 2 and 3, two recta ngular beam eleme nts with 20-i nch wide sect

53、i ons are used betwee n the joist and the Iongitudinal beams of axes 2 and 3 to model the slab in the Iongitudinal directi on. To model the behavior of the slab in the tran sverse directio n, equally spaced parallel beams with 20-i nch wide recta ngular sect ions are used. There is a differe nee bet

54、ween the shear flow in the slab and that in the beam elements with rectangular secti ons modeli ng the slab. Because of this, the torsi onal stiff ness is set equal to on e-half of that of the gross sect ions9.The buildi ng had in fill walls on 2nd, 4th, 5th and 6th floors on the spa ndrel beams wit

55、h some ope nings (i.e. win dows and doors). As men ti oned before and as part of the demoliti on procedure, the in fill walls in the 1st and 3rd floors were removed before the test. The in fill walls were made of hollow clay tiles, which were in good con diti on. The net area of the clay tiles was a

56、bout 1/2 of the gross area. The in-pla ne action of the in fill walls con tributes to the buildi ng stiff ness and stre ngth and affects the build ing resp on se. Ig noring the effects of the in fill walls and exclud ing them in the model would result in un derestimat ing the buildi ng stiff ness an

57、d stre ngth.Usi ng the SAP2000 computer program 8, two types of modeli ng for the in fills are con sidered in this study: one uses two dime nsional shell eleme nts (Model A) and the other uses compressive struts (Model B) as suggested in FEMA356 10 guideli nes.4.1. Model A (in fills modeled by shell

58、 eleme nts)In fill walls are modeled with shell eleme nts. However, the curre nt versi on of the SAP2000 computer program in cludes only lin ear shell eleme nts and cannot acco unt for crack ing. The ten sile stre ngth of the in fill walls is set equal to 26 psi, with a modulus of elasticity of 644 ksi 10. Because the formatio n ofcracks has a sig ni fica nt effect on the stiff ness of the in fill walls, the followi ng iterative procedure is used to acco unt for crack formati on:(1) Assu ming the in fill walls are li